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Full text of "Design of a Low Speed Fan Stage for Noise Suppression"

NASA/CR— 1999-208682 Allison EDR-17923 




Design of a Low Speed Fan Stage 
for Noise Suppression 



W.N. Dalton, D.B. Elliott, and K.L. Nickols 
Allison Engine Company, Indianapolis, Indiana 



Prepared under Contract NAS3-25950 



National Aeronautics and 
Space Administration 

Lewis Research Center 



February 1999 



Available from 



NASA Center for Aerospace Information 
7121 Standard Drive 
Hanover, MD 21076 
Price Code: A24 



National Technical Information Service 

5285 Port Royal Road 

Springfield, VA 22100 

Price Code: A24 



TABLE OF CONTENTS 

Page 
Section Title 

1.0 Introduction 

3 
2.0 Rig Design Features 

5 

3.0 Aerodynamic Design 

3.1 Fan Stage Aerodynamic Design 

311 Baseline Stage Configuration and Vector Diagrams o 



3.1.2 Blade Design 

3.1.3 Baseline Fan Vane Design 

3.1.4 Fan Stage Analysis 

3.1.5 Additional Vane Designs 

3.2 Nacelle Aerodynamic Design 

3.2.1 Inlet Aerodynamic Requirements . 

3.2.2 Aeroline Development 

3.2.3 CFD Analysis 

3.2.4 Aerodynamic Loads 



5 

15 

26 

36 

40 

40 

.49 

.52 

.52 



59 
59 



4.0 Structural Design 

4.1 Rotating Components 

4.1.1 Stress and Deflection Analysis ^ 

4.1.2 Vibration Analysis 

4.2 Static Components 

4.2.1 Stress and Deflection Analysis 

4.2.2 Vibration Analysis 



65 
68 
70 
71 



LIST OF ILLUSTRATIONS 

Figure Title 

Page 

1 Rig mechanical configuration . 

2 Final engine configuration — Task 5 "'"" 6 

3 Rotor exit total pressure profile - 

4 Baseline low noise fan schematic (meridional view) ....."."..."."....". 7 

5 Low noise fan rotor design point profiles o 

6 Low noise fan baseline stator design point profiles ..'.'.' 1 1 ' 

7 Rotor blade geometric parameters "".'.'.'.. 14 

8 Incidence and deviation angles (degrees) " ■, c 



9 Blade near-tip Mach number distribution ."....... 

10 "Blade pitch section Mach number distribution .".....'."."""...... 

11 Blade near-hub Mach number distribution 

12 Empirical modifications to the rotor deviation profile !.!......".......".""." 19 

13 Rotor passage velocity vectors 



19 
20 
21 



25 
26 
27 
28 

29 

30 

31 

32 
33 
34 
35 

36 

37 
38 
39 



.16 
.17 
18 



20 
Daseiine vane arrangement 

15 



14 Baseline vane arrangement 



, 21 

Baseline stator incidence, deviation, and throat margin 22 

16 Baseline stator near-tip Mach number distribution ..."'"."."'. 23 

17 Baseline stator midspan Mach number distribution ,..."."."....."........ 24 

}® Baseline stator near-hub Mach number distribution ! ..."."."."."."."......... 25 

Predicted rotor-only performance map ""'""".'. 27 

Effect of throttling on rotor pressure rise and loss at design speed 28 

Effect of throttling on rotor Mach number and air angles at design speed'.'.'" 29 

2-1 Effect of throttling on blade surface Mach number — near-tip section "30 

23 Effect of throttling on blade passage Mach number - near-tip section ""31 

& Effect of throttling on blade surface Mach number — midspan section '32 

^ ecto f throttlin 8 onblade P assa ge Mach number -midspan section ""33 

26 Effect of throttling on blade surface Mach number -near-hub section... 34 

m Effect of throttling on blade passage Mach number — near-hub section 35 

Rotor blade Mach number distribution at simulated takeoff speed — near-tip 

section K 

; 37 

Rotor blade Mach number distribution at simulated tateoff speed — midspan 

section r 

Rotor blade Mach number distribution at simulated taleoff spe^d 1 — near-hub 

section 39 

Swept vane design point incidence, deviation, throat m argm"a^d Mach niim" 

ber profiles 41 

32 Swept vane design point Mach number distribution^ - near-tip section." "42 

^ Swept vane design point Mach number distributions - midspan section 43 

34 Swept vane design point Mach number distributions - near-hub section 44 

Swept/leaned vane design point incidence, deviation, t iroat margin, and Mach 

number profiles 45 

Swept/leaned vane design point flowfield — near-tip s. Ktion .'.".'.'.".""."".".".' 46 

Swept/leaned vane design point flowfield — midspan ; ection ...".." 47 

Swept/leaned vane design point flowfield — near-hub ;ection Z~.Z1 48 

High bypass ducted propeller drive rig — nacelle layout Rn 

40 Nacelle aerolines „ 

51 



Figure 



LIST OF ILLUSTRATIONS (cont) 
Title Pa S e 



41 PMARC panels 53 

42 Acoustical HBPR nacelle — baseline Mach number distribution 54 

43 Acoustical HBPR nacelle — baseline pressure distribution 55 

44 Acoustical HBPR nacelle — baseline skin friction coefficient 56 

45 Material properties of Ti6-4 (AMS 4928) 62 

46 Material properties of 17-4PH (AMS 5643): H 1100 annealed 62 

47 Low cycle fatigue strength for AMS 4928 (Ti 4-6 STAN forged) at 78°F 64 

48 Low cycle fatigue strength of AMS 5659 (17-4PH) at 70°F 65 

49 Predicted rotor radial deflections 6 ^ 

50 Campbell diagram of blade 67 

51 Goodman diagram of blade 69 

52 Campbell diagram for baseline vane in acoustic testing setup — assembly 

modes '* 

53 Campbell diagram for aft vane in acoustic testing setup — assembly modes 76 

54 Campbell diagram for swept vane in acoustic testing setup — assembly modes 77 

55 Campbell diagram for swept and leaned vane in acoustic testing setup — 

assembly modes '° 

56 Blade track radial deflection versus fan unbalance — pitch mode of acoustic 

testing setup ' 9 

57 Campbell diagram for baseline vane in performance calibration setup — 

assembly mode °*- 

58 Campbell diagram for aft vane in performance calibration setup — 

assembly mode 82 

59 Campbell diagram for swept vane in performance calibration setup — 

assembly mode °3 

60 Campbell diagram for swept and leaned vane in performance calibration setup 

— assembly mode °4 

61 Blade track radial deflection versus fan unbalance — for and aft mode of per- 

formance calibration setup °5 

62 Campbell diagram for baseline and aft vanes — airfoil modes 86 

63 Campbell diagram for swept vane — airfoil modes 87 

64 Campbell diagram for swept and leaned vanes — airfoil modes 88 

65 Goodman diagram for baseline vane 89 

66 Goodman diagram for aft vane 90 

67 Goodman diagram for swept vane 90 

68 Goodman diagram for swept and leaned vane 91 



in 



LIST OF TABLES 
Table Tide Page 

I Flowpath coordinates for LNFB 8 

H Nacelle aerodynamic loads at 15 degree angle of attack 57 

III NASA 22-in. fan rig structural audit checklist — fan blade 60 

IV NASA 22-in fan rig structural audit checklist — fan disk 61 

V Structural audit of static components 72 

VI NASA scaled fan rig nacelle vane static stress sumrr aries 73 

VII NASA scaled fan rig nacelle blade track deflection summary 74 

VIII Flutter parameter vane configurations 89 



IV 



SUMMARY 



This report describes the design of a low tip speed, moderate pressure rise fan stage for demonstration of 
noise reduction concepts. The fan rotor is a fixed-pitch configuration delivering a design pressure ratio of 
1 378 at a specific flow of 43.1 lbm/sec/ft*. Four exit stator configurations were provided to demonstrate 
the effectiveness of circumferential and axial sweep in reducing rotor-stator interaction tone noise. The 
fan stage design was combined with an axisymmetric inlet, conical convergent nozzle, and nacelle to form 
a powered fan nacelle subscale model. This model has a 22-inch cylindrical flow path and [employs .a ro- 
tor with a 0.30 hub-to-tip radius ratio. The design is fully compatible with an existing NASA force bal- 
ance and rig drive system. 

The stage aerodynamic and structural design is described in detail. Three-dimensional (3-D) computa- 
tional fluid dynamics (CFD) tools were used to define optimum airfoil sections for both the rotor and 
stators. A fan tone noise predictive system developed by Pratt & Whitney under contract to NASA was 
used to determine the acoustic characteristics of the various stator configurations. Parameters varied in- 
cluded rotor-to-stator spacing and vane leading edge sweep. The structural analysis of the rotor and sta- 
tor are described herein. An integral blade and disk configuration was selected for the rotor Analyse 
confirmed adequate low cycle fatigue life, vibratory endurance strength, and aeroelastic suitability. A 
unique load carrying stator arrangement was selected to minimize generation of tonal noise due to 
sources other than rotor-stator interaction. Analysis of all static structural components demonstrated 
adequate strength, fatigue life, and vibratory characteristics. 



1.0 INTRODUCTION 



Since the late 1960s, there has been a continuous effort to lower community noise levels resulting from 
aircraft terminal operations. Current high bypass ratio engine technology is sufficient to allow certifica- 
tion of aircraft to Stage 3 of Federal Aviation Regulation (FAR) Part 36. As part of the natural evolution- 
ary process, consideration of a reduced certification level is underway. In order to accommodate a 
g7owth plan, major reductions in propulsion system generated noise wiU be require d te new ««* aft/ 
Line combinations to be certified to this more stringent noise standard. Under Task 5 of contract NAS3- 
25950 Allison Engine Company studied the engine component noise reductions required tc -produce a 
fropul^systernTor a twin engine aircraft producing certification levels 10 decibels (dB). below the cur- 
rent FAR 36 Stage 3 requirement. Early results of this study indicated a strong acoustic advantage m 
movmg from a conventional six bypass ratio turbofan cycle to an ultrahigh ^-^^^1^ * 
low prLure ratio, low tip speed fan. However, cycle changes alone were not sufficient to P^^ 
ver levels 10 dB below stage 3. Additional reductions required identification of innovative strategies for 
Serine the strength of dominant noise sources. Flyover time histories of perceived noise level produced 
unTer ^ Tcontiacf indicated the predominant noise source was the fan during berth the takeoff and ap- 
CachTgments of flight. Noise reduction studies based on this result identified bypass vane sweep as a 
rotentiaS^ffective approach for reducing the pure tone portion of the fan noise held. Based on the re- 
sults of these studies, a fan rig test program was proposed to the National A^^^^^V. 
ministration (NASA) to demonstrate this concept. As a result of this proposal, a 22-mch diameter smgle 
^ ^demonstrator has been designed. This report documents the aerodynamic and structural design 
of this stage. 



2.0 RIG DESIGN FEATURES 



The rig mechanical arrangement evolved from a set of requirements developed to meet the program 
technical objectives and to satisfy facility and operational needs. Specifically these requirements were: 

. the rig must be compatible with existing NASA drive system 
. no flow-path obstructions except rotor and stator allowable 
. provisions must be made for multiple vane configurations 

. vane configuration^ changes must be accomplished in the wind tunnel and not require removal of 
fan rotor 

The final configuration, shown in Figure 1 , meets all design objectives and is fully compatible wiflv the 
NASA drive rig. Based on acoustic analysis, four vane configurations will be tested. Removal of addi- 
tional flow-path obstructions was required to isolate, as fully as possible, the acoustic impact of *e vane 
geometry changes. As a result of this requirement, the stator must carry not only its nonnal aerodynamic 
SS also any nacelle generated loads. To accomplish this and allow vane changes without an rotor 
removal, the vanes have been designed as a segmented ring with the airfoils providing a load ^be- 
tween flange rings on the inner and outer diameters. Loads are passed from the vane ring to a backbone 
™ort though I single shear pin and three radial fasteners in each segment. All outer flow-path pieces 
aft of the rotor are spht axially to allow quick access to the vane fasteners for removal. Multiple attach- 
ment planes are provided to accommodate the four vane configurations to be tested. No provisions have 
been included for either a core rotor or a separate core flow stream. 




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3.0 AERODYNAMIC DESIGN 
3.1 FAN STAGE AERODYNAMIC DESIGN 
3.1.1 Baseline Stage Configuration And Vector Diagrams 

The aerodynamic design point for the fan stage, as established during cycle optimization studies con- 
ducted during Task 5, is: 

Tip Speed = 1000 ft/sec 

Stage Pressure Ratio = 1.362 

W Ve/8 A = 43.1 lbm/sec/ft 2 

As shown in Figure 2, the final engine configuration of Task 5 employed a booster stage on the fan shaft 
to provide the required supercharging for the core compressor. Early in the rig design, it was decided 
both the booster stage and the core flow bifurcation would be eliminated. This produced two benefits. 
The first was a significant reduction in mechanical complexity, resulting in reduced fabrication costs. 1 he 
second was the removal of additional noise sources, allowing a clear identification of the acoustic benefit 
of vane geometry variations. As a result of the very high bypass ratio cycle selected in the Task 5 engine 
study, a strong radial rotor exit total pressure gradient exists (Figure 3). This profile is also present in the 
rig design The rig stage design pressure ratio was selected as the mass average of the 1.38 bypass and 
1 21 core pressure of the original engine design, allowing for some loss through the rig stator. A sche- 
matic cross section of the baseline rig configuration is shown in Figure 4. As can be seen, a cylindrical 
outer flow-path contour was maintained through the stator exit. The requisite area ruling through the 
stage is introduced through the hub flow path as an integral part of the blading design. Curvature was 
used into and through the stator to keep the relatively low momentum fluid coming from the rotor hub 
energized. The rotor-to-stator axial gap is consistent with current Allison fans. Coordinates for the flow 
path of the baseline configuration are presented in Table 1. 

The velocity vector diagrams were generated using the Allison axisymmetric streamline curvature design 
svstem A listing for the aerodynamic design point is included in Appendix A. Some of the blade and 
vane inlet and exit profiles tabulated in Appendix A are plotted in Figures 5 and 6 Also shown are corre- 
sponding profiles from the NASA Stage 53 fan. The comparison is useful since the general character of 
the flow field through the two fans is similar. The NASA Stage 53 fan was designed for the same rotor 
pressure ratio and tip speed; it did not quite pump to design intent, hence, the profiles measured at de- 
sign flow are shown in addition to those that represent design intent. The low noise fan (LNF) rotor is 
designed for a pressure profile of even greater skew and for higher throughflow velocities than found in 
the NASA Stage 53 fan. The rotor inlet is also set for a higher specific flow and lower mlet radius ratio 
(0 30) As a result, the inlet relative Mach number at the tip for the LNF is higher, 1.143, even though tip 
speeds are the same. Greater turning is required across the LNF blade tip but the blade is overall more 
lightly loaded. 

Velocities at stator inlet, although subsonic, are relatively high toward the outer diameter due to the pres- 
sure profile from the rotor and the absence of a splitter. This, together with the thicknesses and camber 
required of these vane sections, made it impossible to design an entirely shock-free stator. Stator loading 
was reduced and performance enhanced by allowing closure of the discharge annulus to a Mach number 
of 0.59 (including blockage). The turning required through the baseline vane row is thus considerably 
less than was required through the NASA 53 fan stator. 

3.1.2 Blade Design 

The fan blade was designed as if destined for a commercial fan application to ensure as much realism was 
incorporated in its geometry as possible. The ultrahigh bypass ratio engine preliminary design cycle was 




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Rotor Pressure Ratio 



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Figure 3. Rotor exit total pressure profile. 



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Axial Coordinate (in.) 



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Figure 4. Baseline low noise fan schematic (meridional view). 



Table I. 
Flow-path coordinates for T.MF* 

Tip contour is a straigh line of radius 11.00000 in. 

HUB 



R 



Spinner nose 




-5.7500 


0.0000 










-5.5000 


0.211325 










-5.0000 


0.5000 


— straight line 








-1.5000 


2.520756 


— segment* 








-1.0000 


2.8000 




Rotor LE 






0.0000 
0.4000 
1.0000 
1.4000 


3.300C 
3.477C 
3.670C 
3.7700 










1.8000 


3.8900 


stack axis = 1.5584 








2.2000 


4.0400 




Rotor TE 






3.2940 
4.0000 
5.0000 
5.6000 
5.8000 
6.0000 
7.0000 


4.4330 
4.6350 
4.8450 
4.9420 
4.9750 
4.9850 
5.1200 




Stator LE 






7.4120 
7.8000 


5.1930 
5.3000 










8.2000 


5.4100 


stack axis = 8.2518 








8.6000 


5.5500 




Stator TE 






9.0930 


5.6000 










10.0000 
11.0000 


5.6000 
5.6000 


— straight line 

— segment 


* R = mZ 


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B 


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Z2^Z^f T 7 Spe6d take ° ff COnditi ° n ' S ° part - S P eed P e *>™ance could be considered. 
Analytically, the blade demonstrates over 16% surge margin at design speed. Leading edge thickness 

£ kSn^ t*t C ° nS1Stent , With Current bird strik * ^teria. TraJmg edge mickLswas seTequal 
to leading ^edge thickness everywhere except near the hub, where a blunter leading edge was employed to 
unprove the hub inlet flow field. Blade chord varies linearly such that the tip is 45% longer tha^Se hub 

£ S"« Z f^S T? ° r^r ***™**<*^ is «*<> *ovm and ranges from 2.75% at thetp 
to 9^42 /o at the hub. The locations of maximum thickness for each secton (not shown) were shifted from a 
uniform 50% chord to improve passage area qualities. Geometric properties are tabulated in AppendkA 



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NASA 53 -Design/ 

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0.50 0.60 0.70 0.80 0.90 1.00 1.10 1.20 

Inlet Relative Mach Number 



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nlet Relative Flow Angle 



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Stator Inlet Meridional Velocity 



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Stator Inlet Row Angle 



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Stator Exit Meridional Velocity 



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90.0 
80.0 
70.0 
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50.0 
40.0 
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Nasa 53 - Design 

O -ONasa 53 - Data 



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Stator Exit Flow Angle 
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Figure 6. Low noise fan baseline stator design point profiles (1 of 2). 



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70.0 
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Stator inlet Mach Number 



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i ■ ■ ■ p i 



— Baseline Design 

I Nasa 53 - Design 

IG- -O Nasa 53 - Data - 

\ 




■o 

1 ' ' I ■ ■ ■ j I ■ ■ ■ ■ N . T . . t . ■ . , 



0^0 0^5 0.30 0.35 0.40 0.45 0.50 

Stator Diffusion Factor 



100.0 
90.0 
60.0 
70.0 

60.0 

c 
to 

°- 50.0 



CO 

If 



40.0 
30.0 
20.0 
10.0 



0.0 



■ ' 1 1 | I I T 1 | 1 1 1 1 [ 1 1 1 | || t 


1 I | 1 T ! 1 | 1 11 f 


_ Baseline Design 1 


l^ \ - 


Nasa 53 - Design | 


* " 


" O -ONasa53-Data 


/ 


1 / - 
1 / . 


/ 


^ /- 


/ 


/ / 


/ / 


1 1 


/ ' 


j 


: / ? 


/ - 


i ' 




f * / 




■ I'V 


- 


/ / 




/ A 




/ ks 




'<*'/ 




1 1 1 1 l/\ 1 1 1 \'\ \ | 1 1 | 1 | | \ 1 L , 


■ 1 i > i i 1 . • ■ . 



0.35 0.40 0.45 0.50 0.55 0.60 0.65 0.70 

Stator Exit Mach Number 



100.0 
90.0 
80.0 
70.0 
60.0 
50.0 
40.0 
30.0 
20.0 - 
10.0 - 



■ | ■ 1 1 1 ; 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 



ji iy i ii ■ I •••• 



'©. 



tiaseline Design 

tlasa 53 - Design 

O- -OHasa 53 - Data 



Sf 



0.0 



1 ■ ■ ■ I " " i ■ ■ " ' ■ Aa. ... i .... i i 



0.60 0.65 0. 0.75 0.80 0.85 0.90 0.95 1.00 

Stag 5 Adiabatic Efficiency 
2 of 2 
TE96-1018 



Figiore 6. Low noise fan baseline stator design poir t profiles (2 of 2). 



12 



Preliminary design of the blading was carried out assuming multiple-circular-arc (MCA) airfoil sections. 
OveX tow tip speed of this fan, an MCA blade was acceptable for studying the effects of changes in 
aspect ratio, maximum thickness, and spanwise chord distributions on surge margin and mechanical in- 
tegrity. The final blade is made up of sections of aerodynamically-optimized meanlines with near- 
sinusoidal thickness distributions. Viscous computational analysis was used extensively to obtain ttie 
desired match of the blade passages with the design intent flow field. The transonic sections were tai- 
lored for the design speed shock structure permitting the largest excursion in flow range to stall with ac- 
ceptable performance. 

The spanwise distribution of incidence angles to which the blade sections were set, shown as *««Jj^« 
in Fieure 8 evolved from several considerations. One was the decision to design to relatively tight throats 
(3.5%throat margin) to favor operating line performance. In the portion of the blade with fupe^omc m- 
et relative flow, another consideration was to observe the first captured Mach wave ruk. This is a rule- 
of-thumb setting a critical incidence off the suction surface at a point halfway between the leadmg ; edge 
and the point of emanation of the first captured Mach wave to a minimum of 1.5 degrees, to ensure flow- 
handling capability. A third consideration involved the meanlines of aU sections which were carefuUy 
shaped to produce acceptable surface Mach number distributions devoid of local peaks or ■spikes. This 
could be done over the outer half of the blade only by straightening the meanlines forward of the throat 
locations and forcing the bulk of the turning aft (Figure 9). Where possible, the subsonic sections were 
taUorecTfor shock-free (design point) operation. Optimum chordwise loading distributions were achieved 
Iv keeping me^nline curvatSewell f omard and closing the leading edge. All this led to incidences con- 
siderably smaller than employed in the design of the NASA Stage 53 rotor. 

The predicted surface distributions of isentropic Machnumber and associated passage Mach number 
contours for the near-tip, pitch, and near-hub sections of the blade are shown in Figures 10, 11, and 12. 
The Ta -tip section was fashioned to produce a single, oblique shock pulled well back into the passage 
and impinging on the suction surface just ahead of the region of greatest curvature. The suction surface 
Mach /umber rises smoothly to a peak of about 1.35. The pitch section, shown in Figure 11, was shaped 
to operate shock free. Maximum thickness was brought forward to the mouth and curvature was±s- 
tributed over a larger portion of the section to flatten the forward portion of the suction surface velocity 
distribution. 

Area-ruling of the hub flow path was an integral part of the design of the near-hub sections. D"«-to 
thickness me hub was found to be quite insensitive to incidence and local meanlme changes. Modifica- 
tion of the hub flow path improved the loading distributions. The intent was to forcethe section Wing 
forward without allowing the hub to overpump (due to greatly increased camber). Several iterations 
were required, with the final outcome shown in Figure 12. 

The rotor deviation angles, shown in Figure 8 were set by augmenting calculated NASA _ 2-D ™le devia- 
tions with the empirically-estimated corrections plotted in Figure 13. These corrections have been estab- 
lished through comparisons of computational and measured results from other Allison <^P^sor 
stages, as well as published reports. The computational results suggest, for the deviation distribution 
chosen, there is sufficient camber in the blading to produce the desired pressure profile. The velocity 
vectors for the near-tip, pitch, and near-hub sections reveal a healthy flow field with no trace of incipient 
separation (Figure 14). 

The static or manufactured blade geometry producing the desired blade shape at design speed was de- 
termined by subtracting the predicted deflection of the blade due to centrifugal and aerodynamic loadmg 



13 



100.0 
90.0 
80.0 
70.0 

60.0 

CO 500 

40.0 
30.0 
20.0 
10.0 |- 



0.0 ' ■ ' ' 




3.00 3.50 4.00 4.50 

Chord (in.) 



5.00 




0.00 0.02 0.04 0.06 0.08 0.10 0.12 

Max Thickness/chord 



c 

CO 
CO 



100.0 
90.0 - 
80.0 - 
70.0 
60.0 
50.0 
40.0 
30.0 
20.0 
10.0 - 



0.0 



LmdngEdgs 

Trailing Edge 




100.0 I i i i j | i i i i | i i ■ i [ ■ i , i | , 



000 0.02 0.04 

Edge Radii (in.) 




0.0 I » i ■ i I ■ i i i I . ■ ■ i . , , . i .x . I 

10.00 15.00 20.00 25.00 30.00 35.00 40.00 

Camber (deg.) 

TE96-1019 



Figure 7. Rotor blade geometric parameters. 



14 



100 r 



c 
re 
a. 




-4-3-2-10123456789 

Incidence and Deviation Angles, degrees TE96-1 020 

Figure 8. Incidence and deviation angles (degrees). 

applied at the design point. These deflections were determined using an Allison proprietary finite ele- 
ment structural analysis procedure. Airfoil sections are defined on planes normal to the stack axis. The 
stack axis is a radial line passing through the center of gravity of each conical section. The leading edge 
shapes are elliptical. The blading opens with speed by as much as 2 deg in stagger at the tip, due mostly 
to flexibility of the leading edge. Associated with this movement in the blade-to-blade view, which 
clearly affects flow handling and pumping capacity, is the radial growth of the tip, with its consequences 
on clearance effects. 

3.1.3 Baseline Fan Vane Design 

A view of the baseline stator design, fan configuration No. 1 (FC1), is shown in Figure 14. Unlike the ro- 
tor the stators are unique to the 22-in. NASA rig vehicle because none could be directly scaled-up for use 
in a high bypass turbofan. In an engine, separate stator assemblies would be required for the bypass and 
core flow streams. Neither of these assemblies would necessarily reproduce a section of the rig stators, 
due to the presence of the flow splitter. Nevertheless, the stators are crucial components of the rig tests. 
The baseline stator must deliver the same performance and allow no more noise in the acoustic test vehi- 
cle than would the bypass stator in a representative commercial turbofan. 

The dominant feature of the stator flow field is its nonuniform, high-velocity inlet (Figure 6). The baseline 
stator is relatively lightly loaded and does not have to affect a large amount of turning, so the emphasis 
during its design was on minimizing total pressure loss. As a result, the vane design process primarily 
involved the selection of an incidence distribution. For any given incidence, neither meanline shape 
maximum thickness, nor section thickness distribution had any appreciable effect on performance. There- 
fore, simple double circular arc sections with maximum thickness located just forward of mid-chord were 
employed. 



15 



A, 



Relative Mach 
Number 



1.4 



1.3 



1.2 



1.1 



9 5.8% Span 




12 
I 3 



16 
17 



20 
21 
22 



. I«4£. 10 

. JOE- 1 

.113 

. 1(9 

. 225 

. 212 

. 331 

. 394 

. 451 

. J07 

. ]<3 

. <20 

. <7« 

. 732 

. 719 

. (43 

. 901 

.957 

I. 01 

I. 07 

1.13 

I. It 

1. 24 

I. 30 

I. 35 



JT ^ 



0J 
XI 

S 
5? 

•3 



.a i.o 

a 

o 

a 

CD 
CO 



0.9 



0.8 



0.7 



r 



\ 




I 




I 

\ 



■S 




0.2 



04 0.6 

Fraction of Chord 



0.8 1.0 

TE96-1022 

Figure 9. Blade near-tip Mach number distribution. 



16 



Relative Mach Number 



1.20 




54.3% Span 




1 - 


. I6SE- 01 


2 " 


. 4 3 IE- 1 


3 " 


. 902E-01 


4 • 


. 133 


5 • 


. 110 


6 • 


. 226 


7 


. 271 


> • 


. 316 


9 


. 361 


10 • 


. 406 


1 1 - 


. 431 


12 • 


. 496 


13 - 


. 341 


14 • 


. 316 


13 ■ 


. 632 


16 - 


. 677 


17 . 


. 722 


li • 


. 767 


19 • 


.112 


30 • 


. 137 


21 ■ 


. 902 


22 • 


. 947 


23 • 


. 992 


24 . 


1. 04 


23 ■ 


1. 01 



0.40 



0.0 



0.2 0.4 0.6 0.8 1.0 

Fraction of Chord 

TE96-1023 

Figure 10. Blade pitch section Mach number distribution. 



17 




Relative Mach Number, 5.9% Span 



1.0 r- 




s 

O 

CO 



a 
o 

is 

c 

a) 



0.8 - 



0.7 



0.6 



0.5 - 



0.4 



0.3 



s 

< 

7 
8 

» 

10 
I 1 
12 
1 3 
14 
1} 
It 
1 7 
II 
19 
20 

n 



. 293E-09 
. 374E-01 
. 74IE-01 
.112 
. 150 
. 117 
. 234 
. 262 
. 299 
. 337 
. 3T4 
.412 
. 449 
. 4I« 
. 324 
. 361 
. 399 
. 63« 
. «73 
.711 
.III 



f 





0.2 



0.4 0.6 

Fraction of Chord 



0.8 l.o 

TE96-1024 



Figure 11. Blade near-hub Mach number distribution. 



18 



100 1— 



90 



80 



70 



60 



c 

CO 

& 50h 



40 - 



30 



20 



10 




NASA Stage 53 Rotor 



-1 



1 2 3 

"Delta Deviation" Angle, degrees 



5 6 

TE96-1025A-2 



Figure 12. Empirical modifications to the rotor deviation profile. 



19 



\ 



^ 



\ 
\ 



Near Midspan - 54 



Near Tip - 95.8% Span 





\ 



\ 

\ 

\ 



\ 
\ 

\ 
\ 

\ 
\ 

N . 



Near Hub -5.9% Span 




TE96-1026 



Figure 13. Rotor passage velocity vectors. 



20 




TE96-1027 



Figure 14. Baseline vane arrangement. 

^ge^e loading. All attempts to combine the two types radially forced the outer sections toward A 
type distributions. 

c- 11 u~c<™ f nr fpi The desien offered reduced suction surface Mach number 

was adjusted to remove all swirl as would be required of a bypass stator. 

The surface isentropic Mach number distributions and associated passage Mach ^ e ;^^ a ; s e ec . 

balance the loading distributions of the near-hub sections. 

The mechanical properties of the baseline vane are tabulated in Appendix A^ These F ^^ 16 ' 
SneTuVdesignmg the alternative vanes. Most have constant spanwise distributions; e.g. maxunum 
thickness-to-chord is 5% and chord is 1.81 rn. 



21 



i oo r 



80 



c 

CS 

a, 

CO 



60 



40 



Q 



/ 
/ 



®/ 

/ 
/ 
/ 
/ 
/ 
/ 



/ 



J ! L 



Incidence, deg. 



, (A) Closed Leading Edge 

i (B) Open Leading Edge 
/ (Final Design) 




'■ °* l- 08 1.12 1. | 6 

Throat Margin 



1. 2 



ioo r 



c 

CS 

CO 




TE96-1028 



Figiire 15. Baseline stator incidence, deviation, and throat margin. 



22 



1.3 r 



1.2 




Relative Mach Number 




. OOOR«00 
. 445E- 01 
. I90E- 01 
. 134 
. Ill 
. 223 
. 162 
.312 
. 336 
. 401 
. 443 
. 490 
. 334 

. 579 

. 623 

. 6«l 

.712 

. 757 

. 101 

. 146 

. «90 

. 935 

. »79 

1. 02 







V 



I ■ 

> 



95.0% Span 



0.2 



0.4 0.6 

Fraction of Chord 



0.8 



1.0 



TE96-1029 



Figure 16. Baseline stator near-tip Mach number distribution. 



23 




1.1 r- 



1.0 



i 

2 
J 

4 

s 

e 

7 

i 

9 
10 

] 1 
II 

15 
14 
13 
16 
17 
II 
19 
20 
21 
22 
2) 
2< 



OOOE.00 
381E-OI 



Relative Mach Number 



u 

o 

S 

o 

cB 



a 
o 

c 

V 
09 



0.9 



0.8 



r 




0.7 - 



0.6 




0.5 



0.4 



r 



0.3 




0.2 



0.4 0.6 

Fraction of Chord 



0.8 



1.0 
TE96-1030 



Figure 17. Baseline stator midspan Mach nuriber distribution. 



24 





0.80 



|~ Relative Mach Number, 9.5% Span 



VALUES 




1 » 


0005*00 


2 


3066- 


01 


3 - 


6I2E 


01 


4 


9I8E 


01 


5 - 


122 




6 - 


153 




^ • 


. IK 




I ■ 


.21* 




9 - 


. 245 




10 - 


. 276 




i 1 • 


. 306 




1 2 - 


. 337 




13 - 


. 367 




14 - 


. 391 




IS - 


. 429 




16 ■ 


.459 




17 . 


. 490 




18 - 


. 520 




19 - 


. 551 




20 - 


. 512 




21 ■ 


. 612 




22 - 


. 643 




23 - 


. 674 




2( - 


. 7»« 






0.40 



0.2 



0.4 0.6 

Fraction of Chord 



0.8 



1.0 
TE96-1031 



Figure 18. Baseline stator near-hub Mach number distribution. 



25 



3.1.4 Fan Stage Analysis 

3.1.4.1 Predicted Map 



The predicted 100% and 85% speedline characteristics for the low noise fan (LOT) rotor are a composite of 
analytical and empirical considerations (Figure 19). The shaded circles in the figure e P re7entSy S f 
results at various backpressures for 100% and 85% corrected speed. No attempts made ^ ^lud the 
untwist charactenstacs of the blade with speed or throttling. To model the indfcated ^Zk££ 
point, the code was run to an "equivalent" design point just over 1% higher in flow and preSuTratio 
This was done in hght of prior experience with the code to be explained in section 3.1.4.2 below T^e 
background speedlmes result from scaling the experimentally-derived map of the NASA Stage 53 rotor to 

NAsTsTd'* dCSlgn P 0lnt "* are f duded for ref — to trends only. The spee^e^ell^Z 
NASA 53 data roughly corresponds to the computationally predicted behavior of the current S T The 
design mtent surge margin of 15% was obtained. The associated contours of predicted effSencyTe also 
hown with the NASA Stage 53 rotor data, scaled for flow, in the background*^ These dltaZZ7ZredTi- 
ficult to assess. The computational procedure, at least for high speed machines, typicaUv predTctTeffi 
ciencies 2 to 3 l points higher than are actually attained; this has been assumed a fu^tionXmputaLal 

^SSTSTT* *k C ° d !, With SUffidently d6nSe ^ ds t0 a :curatel y re P-duce prSrag Te 
f^JT I^ZT' ** P f? Cted effkienCy haS been modified to bett « & the available data T 
general, the modified efficiency follows the trends predicted by the code, but reduced at the design con- 
dition to correspond to the value obtained from the axisymmerric streamline curvature proceXflddi 



3.1.4.2 Off-Design Performance 



The LOT rotor, though part of a research vehicle to be built for acous tics testing, was designed to stan- 
dards allowing it to be scaled-up directly for use in a large turbofan rngine fS that reasS an effl* w« 
made to ensure the blade would also demonstrate good off-design perfSmance it Z^Zd^onT 
tZZTl % nea Tf^ d 3t "* -^^ Condition at bo* 100% and S^LTi^tf^ 
through 27 show how the LOT rotor is expected to throttle at design speed. »gureszu 

^p% C ^T^t ^ ? C *? P reSSUre ""* l0SS P rofiles of *" rotor with ^"^g are shown in Fig- 
ure 20 The long dashed line labeled "ADP-BD76" is the design inten profile from me axisvnuTtS 

ST^r C °u de - ** *** 0th6r ^ are * e P rofil -predicted by th'nunS solu- 
tion at the tfuree points along the design speedline highlighted intte :nap of Figure 19. The^FD solution 

SSSK?? md ' CateS 3 f ° nger hub * nd a WGaker ti P *« *** s * d «velop in reanty, so *e S 
labeled ADP-Dawes" was selected as the one to use for the detailed design of the blade Here a*aT*e 
analysis of the NASA Stage 53 rotor flow field proved useful. The differences between ieBDVeSf 
computational profiles for that machine were considered in establishing the LOT des^ proffle ^e m - 

^ r^t T aU Pre T e Pr0fileS Mkate pUm P m S at ** hub <"»** w °uld ddiver the core flTw 
m the tuAofan), remains unchanged while the bypass portion of the blade, from 20% span to the tip 

S a ZlTT y W1 ^ US - I" SSeS klCreaSe Wi * throttIin 8 ta a consistent — - excepturi- 
ously at tiie near-tip near stall where they apparently decrease. The c hanges in throughflow velocities 
are reflected in the profiles of inlet relative Mach number and air ang] es (Figure 21). aTS^SST 

ctLt^ ^^rV^ ^ fraCti ° n ° f fl ° W throu 8h the hub decrees. The h^Se^L- 
creases 5-6 degrees while the tip increases only 2-3 degrees. The disci arge air angles remain little 
changed over virtually the entire blade span, another indication there is LffidenfcamSTiSjlade and 
the turning can be sustained without a breakdown in the flow field right up to stall. 

The predicted changes in surface Mach number distributions and pass age Mach number contours for the 
near-tip, pitch, and near-hub sections with throttling are shown in Fig ares 22 through 27. Mosi "noticeable 



26 



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29 



95.8% Span 



Unthrottled 




0.2 



0-4 0.6 

Blade M/M 



0-8 i.o 

TE96-1 035-2 



Figure 22. Effect of throttling on blade surface Mach 



numler — near tip section. 



30 



Near Stall 



Design Point 





ii 

19 




4E- 10 
JE-OI 



Unthrottled 



JE- 10 
06-01 



TE96-1036 



Figure 23. Effect of throttling on blade passage Mach number — near-tip section. 



31 



54.3% Span 



Unthrottled 




0.4 0.6 

Blade M/M 



0.8 1.0 

TE96-1 037-2 



Figure 24. Effect of throttling on blade surface Mach number — midspan section. 



32 



I) 




Design Point 



Unthrottled 




VALUES 

. . I 69E-0I 
. .546E-0I 

• . 109 
■ . 164 
. . 219 
> .273 
. . 331 
. . 312 

• . 437 
• . 492 

346 

601 

656 

710 

763 
. 120 
. «74 
. 929 
. 9t3 
I. 04 

I. 09 

1.15 

23 - 1. 20 

24 ■ 1. 26 
2J - 1.31 



11- 

12 ■ 

13 - 

14 - 

15 - 

16 - 

17 . 
It - 



VALUES 

. . 16IE-0I 
. . 4 S 1 E- 1 

I . .902E-0I 

I . .133 

I . . ISO 

S . .226 

I . . 271 

I • . 316 
t • . 361 
10 - . 406 

II . .451 

12 - .496 

13 - .541 

14 • . 316 

15 • . 632 

16 - .677 

17 . .722 
It .767 
19 • .112 

. 157 
. 902 
.947 
. 992 
I. 04 
1. Ot 



20 
21 
22 



TE96-1038 



VALUES 


■ 


224E-09 


1 


3 36E-0I 


J - 


1 11 


I ■ 


167 


5 • 


222 


6 


271 


7 « 


334 


1 • 


319 


9 


. 443 


10 ■ 


. 301 


1 1 - 


. 556 


12 • 


.612 


13 • 


. 667 


14 . 


. 723 


15 ■ 


. 779 


16 - 


.134 


17 • 


. 190 


It • 


. 943 


19 - 


1. 00 


20 - 


1. 06 


31 • 


111 


22 . 


1. 17 


23 - 


1. 22 


24 - 


1. 21 


23 - 


1. 33 



22 



Figure 25. Effect of throttling on blade passage Mach number • 



- midspan section. 



33 



1.2 i— 



5.9% Span 




0.2 



0.4 0.6 

Blade M/M 



0.8 1.0 

TE96-1 039-2 



Figure 26. Effect of throttling on blade surface Mach 



number — near-hub section. 



34 




Near Stall 



. 290E- 09 

. 3 24E- 01 

. 649E 01 

. 973E-0I 

. 1 30 

. 162 

. 195 

. 217 

. 260 

. 292 

. 324 

. 337 

. 319 

. 422 

. 454 

. 4t7 

.519 

. 552 

. 514 

. 616 

,411 



Relative Mach Number, 5.9% Span 



Design Point 




VALUES 




1 ■ 


293E 


09 


2 


J74E 


01 


} 


74IE 


01 


4 - 


1 1 2 




5 - 


130 




6 ■ 


117 




7 > 


224 




1 • 


162 




9 ■ 


299 




10 • 
1 1 ■ 


337 

374 




1 2 • 


4 1 2 




1 3 • 


449 




1 4 ■ 


416 




IS ■ 


324 




1 6 ■ 


361 




1 7 • 


599 




It • 


636 




19 - 


673 




20 • 


71 1 




:i ■ 


III 





r 



1 1 



r 



1 5 



Unthrottled 




. 295E- 09 
. 452E- 01 
. 904E- 01 
. 136 
.III 
. 226 
. 271 
. 317 
. 362 
. 407 
. 452 
. 497 
. 543 
. SIS 
. 633 
. 678 
. 724 
. 769 
.114 
. 859 
111* 



TE96-1040 



Figure 27. Effect of throttling on blade passage Mach number — near-hub section. 



35 



is the movement of the shock near the tip. It migrates from within the passage, as an over-expansion 
normal shock (probably a reflection from the suction surface of a veiy weak leading edge oblique shock) 
to a strong, started oblique leading edge shock to an unstarted, though still stable, normal position to a 
final, (not shown) unstable interaction with the pressure side bow waves from the neighboring blade. It is 
the ability of 3-D codes to reproduce shock system geometries and reveal the effects on performance of 
shock structures that make them such powerful design tools. Note the problem of suction side peakiness, 
discussed earlier, cannot be avoided in the unthrottled condition. Tie shock is far enough aft that it im- 
pinges on the blade in the region of greatest curvature. A vestige of the near-tip flow field can still be 
seen in the pitch passages, though it was possible to design the pitch section for shock-free operation at 
the design point. The hub experiences the largest change in flow level and not surprisingly, becomes the 
pinch point with decreasing backpressure. The surface Mach number distributions illustrate the rationale 
for the selection of incidences discussed earlier. The progression from negative to design to larger inci- 
dences with throttling is apparent. 

Once a blade shape acceptable at design speed was defined, it was analyzed at 85% speed, which was 
defined as takeoff speed in the ultrahigh bypass engine cycle. At thrs speed, the blade tip inlet runs to just 
under Mach 1.0. Obviously, the nominal operating line condition at this speed is particularly important 
from a noise production standpoint. The surface Mach number distributions and passage Mach number 
contours predicted for the near-tip, pitch, and near-hub sections at the takeoff point are shown in Figures 
28, 29, and 30. Notably, even the outermost section operates shock free. 

Incidence levels are uniformly higher than at design speed. The near-tip and pitch sections also exhibit a 
pronounced reacceleration bump in their suction surface Mach number distributions. This is produced by 
the large local curvature in each section, discussed earlier, that is in nirn one consequence of designing to 
relatively tight throat margins. 

3.1.5 Additional Vane Designs 

The NASA test plan calls for the acoustic evaluation of four distinct configurations. Each is characterized 
by a different stator; the rotor design described earlier is common to all. The baseline fan includes the 
radial vane already described. The second configuration of the fan, designated FC2 and shown in Figure 
lb, results from repositioning the baseline stator further downstream and increases the rotor-to-stator ax- 
ial gap. Although the vanes are placed in a slightly different flow field, as modeled in Appendix B, the 
stator assembly itself remains unchanged. The third and fourth fan configurations, however, necessitated 
the design of two new stator vanes and associated flow-path modifications. For fan configuration No. 3 
(FC3) the stator of the baseline fan is replaced with a vane whose leading edge lies at a 30 degree angle 
from vertical. The fourth fan, FC4, replaces this stator with another made up of vanes that are both swept 
and leaned. These latter two stator designs are described below. 

3.1.5.1 Axially Swept Vane Design 

From an acoustic study conducted by NASA, it was determined that among a candidate set of purely 
swept shapes, a vane swept 30 degrees aft offered the best potential for noise reduction. A stator with 
this amount of sweep was designed so the radial vane stator of FC1 could be replaced, requiring only one 
additional spoolpiece to recomplete the outer casing. That the extra spoolpiece was required in any case 
proved fortuitous since it was found during design of the swept vane that the outer flow path could not 
be kept of constant radius. A meridional view of the swept vane fan, FC3, is shown in Figure lc. Not 
only is the vane highly swept but the casing forward of and through the vane includes a substantial 
bulge. The incorporations of sweep so increased throughflow veloc tries that changes in airfoil sections 
alone were not enough to produce satisfactory outboard vane passage designs; careful area-ruling of the 
flow-path annulus had to be considered at the same time. Several casing and hub wall contours were 
analyzed to optimize the final flow-path geometry. 



36 



u 
J5 



•s 

OS 

'S. 

o 

c 

tt> 



1.2 r 

" ft 

0.9 - 

0.8 

0.7 

0.6 

0.5 



.J"""* 



■«-■-■' 






■ 

\ 



'«■■■ 





0.2 0.4 0.6 

Fraction of chord 



0.8 



1.0 




VALUES 




1 


. 1 77E- 


09 




. 465E 


01 




. 930E 


01 




. 140 






.116 






. 233 






. 279 






. 326 




9 


. 372 




10 • 
1 1 


.419 
. . 465 




1 2 ■ 


■ .5 12 




1 3 


■ .558 




1 4 


■ . 605 




15 


■ .65 1 




16 


■ . 698 




17 


> .744 




1 < 


. .79 1 




1 9 


■ . «37 




20 


• .114 




21 


. . 930 




22 


. . 977 




23 


. 1.02 




24 


. 1.07 




25 


- 1.12 





TE96-1041 



Figure 28. Rotor blade Mach number distribution at simulated takeoff speed — near-tip section. 



37 



0.95 



0.90 



0.85 



0.80 



, 



S\ 



\ 



\ 



0) 



2 0-70 H 

« 

S 

•§, 0.65 
2 

| 

►5 0.60 



0.55 
0.50 
0.45 



0.40 



\ 



■ 




J_ 



I 



0.4 0.6 

Fraction of chord 




21 
2} 
24 
25 



0.8 



I37E-09 
314 E- 01 

76IE-01 
I 13 
134 
1 92 

30 

«9 

07 

43 

14 

22 



1.0 



" TE96-1042 



Figure 29. Rotor blade Mach number distribution at simulated takeoff speed — midspan section. 



38 



l.o r 



0.9 |- 




■■■■■-■-«., 



::* 



_L 



_1_ 



0.2 0.4 0.6 

Fraction of chord 



0.8 



1.0 





it 



. 641E- 10 
. 323E-01 
. 646E-01 
. 970E-0I 
. 129 
. 1C1 
. I 94 
. 226 
. 239 
. 291 
. 323 

. 356 

. 311 

. 420 

. 453 

. 413 

.511 

. 549 

. 512 

.614 
1)1 



TE96-1043 



Figure 30. Rotor blade Mach number distribution at simulated takeoff speed — near-hub section. 



39 



The design objective was a swept stator with the same kind of velooty distributions over the vane sur- 
faces as were obtained with the baseline vane. All of the physical properties of the baseline vane i e 
^TTu 311 gC thicknesses ' <****> location °f maximum thickness, etc, were preserved in both this 
and the following swept vane designs. Double circular arc sections were employed as before although 

W 11 *n T^J i 6 ^ 00 T ? S f CTeaSed - SinCe ** ° Uter diame :er bul S e not onl y redu «d the level 
but also flattened the shape of the throughflow velocity profile, incidences were adjusted accordingly 
Profiles of these parameters are shown in Figure 31 compared with those for the radial vane, at the respec- 
tive vane edges. The surface isentropic Mach number distributions and associated passage Mach number 
contours are shown in Figures 32, 33, and 34 for the near-tip, midspan, and near-hub sections shown for 
^endKc me V3ne ' S amng FC3 C ° nditionS at the adynamic design point is included in Ap- 

3.1.5.2 Swept and tangentially Tilted Vane Design 

The NASA acoustic study referred to previously indicates a potential for further noise reduction by add- 
ing lean tangential bit) to a swept vane. The study suggests a vane leaned 30 degrees suction-side down 
(toward the I.D.) with the lean, like the sweep, incorporated so the vane edges remain straight (viewed 
along engine centerlme) offers the largest benefit. The FC4 vane was designed for this degree of stack 
axis lean. The fined geometry is shown in Figure Id. Noticeably absent is the large bulge in the O.D. flow 
path required in FC3. Referring to Figure 35, it can be observed that vane lean increases the flow block 
age, producing a proportional increase in throughflow velocity, but tends to reduce the migration of flow 
toward the outer flow path compared to the simple swept design. As a result, flow-path contouring up- 
stream of the leading edge is not required. Deviation shows a strong sensitivity to loading. In the out- 
board sections, increased loading produces an increase in inlet-to-dis-iiarge velocity rations a result de- 
viation angle increases. For the inboard sections, increased loading r, suits in a decrease in section ve'loc- 
lty ratio; as a result, the deviation angle decreases. 

As for FC3, the design objective for the swept and leaned vane was to reproduce velocity distributions 
over its surfaces as much like those obtained for the radial vane as possible. Section incidences were 
adapted to help achieve this. The resultant distributions for the usual three sample sections are shown in 
AppeTdk D § aUinS FC4 C ° nditions at *** adynamic design point is included in 

3.2 NACELLE AERODYNAMIC DESIGN 

This nacelle design was developed to meet the basic operational requirements of an isolated nacelle con- 
figuration for subsonic/transonic application having an advanced turjofan inlet and a separate flow ex- 
haust system. Since the test vehicle includes no provisions for a separate core flowstream, the primary or 
core nozzle was truncated and replaced by the propulsion rig metering strut housing the powered drive 
Thus, the inlet flow equals the fan nozzle exit flow, unlike a turbofan flight nacelle where the inlet flow ' 
splits mto the fan and the core flow and exits separately from the two sxhaust nozzles. 

3.2.1. Inlet Aerodynamic Requirements 

Inlet Dimensions 

Both inlet and exhaust systems are sized for maximum inlet corrected flow of 102.78 lb. At this flow rate 
the mlet throat area is designed for maximum average Mach number et the throat (M m ); to be equal to or 
below 0.75. At this flow condition, the fan operates at maximum speci fie flow of 43.5 Ib/ftZ This value is 
consistent with current Allison Engine Company fan design criteria. 



40 





uBdg % 





tredg' 




Oh 

1-1 

1 

3 
C 

A 
u 



C 

.s 

60 
re 

s 

o 



c 
o 



> 

CJ 

c 

4> 
TS 

'u 
C 

.s 

o 

-I 

<n 

CJ 
TJ 

01 

C 

> 

a, 



CO 

3 



xreds% 



41 



Relative Mach Number 



u 
£ 

1 



a 
o 

u 

C 




VALUES 



I 


. OOOEfOO 


2 


. 45IE-01 


3 . 


- 901E-0I 


4 . 


. 135 


5 


. 110 


6 


. 225 


7 . 


. 270 


1 ■ 


.315 


9 ■ 


. 361 


10 - 


. 406 


1 1 . 


. 431 


12 ■ 


. 496 


13 . 


. 541 


14 - 


. 516 


IS - 


. 631 


1« • 


. 676 


17 • 


. 721 


11 • 


766 


19 • 


111 


20 ■ 


«56 


21 • 


901 


22 . 


946 


23 . 


991 


2* . 


. 04 



0.2 



0-4 o.i> 

Fraction of chord 



0.8 



J 
1.0 

TE96-1045 



Figure 32. Swept vane design point Mach number disl ributions - 



near-tip section. 



42 



49.8% Span 



VALUES 


I - • 


OOOE»00 


2 - ■ 


372E-01 


3 - - 


744E- 01 


* ■ . 


112 


5 - 


149 


6 • 


116 


7 ■ 


223 


1 - 


260 


9 - 


29* 


10 - 


.335 


1 1 « 


. 372 


12 • 


. 409 


11 - 


. 446 


14 • 


. 414 


15 - 


. 521 


16 - 


.55* 


17 • 


.595 


11 • 


. 632 


19 - 


. 670 


20 - 


. 707 


21 - 


. 744 


22 - 


. 7*1 


23 - 


. 11* 


24 - 


. *56 


25 • 


. *93 




0.2 



0.4 0.6 

Fraction of chord 



0.8 



TE96-1046 



Figure 33. Swept vane 



design point Mach number distributions - midspan section. 



43 



9.5% Span 




VALUES 




1 


. 000E 


► 00 


2 - 


. 327E 


01 


3 ■ 


. 654E 


01 


4 . 


. »<0E 


01 


5 - 


. 131 




6 m 


. 163 




7 > 


. 196 




t . 


. 229 




9 • 


. 261 




10 - 


. 294 




1 1 - 


. 327 




12 • 


. 359 




13 • 


. 392 




14 > 


.425 




15 - 


. 451 




1< • 


. 490 




17 > 


523 




It » 


556 




19 • 


Si( 




20 - 


621 




21 - 


654 




2« . 


• II 





Relative Mach Number 



0.80 r- 



0.76 



0.72 



-Q 
S 

I 

j: 0.64 



o 

hi 

C 

0J 



0.56 



0.52 



0.48 



0.44 




0.2 0.4 o.6 

Fraction of chord 



0.8 



1.0 



TE96-1047 



Figure 34. Swept vane design point Mach 



number distributes — near-hub 



section. 



44 






00 

o 



H 



C 

'5b 






e 
o 



> 
o 

Q 



O 

c 



1 

3 
C 

u 



T3 

C 

(8 



60 
s- 
(0 



« 
O 
u 

C 

o 



> 

01 

-a 

(J 01 

QJ »-i-t 

"S £ 

"0 Oh 



.S 

ex 

c 

Ol 
T3 
Ol 

> 

o> 
c 
« 
o> 



01 
U) 

Lri 

ro 

01 

ii 

3 
60 



uBds% 



UBdg % 



45 



\ 



94.9% Span 




Relative Mach Number 



1.1 



1.0 - 



-a 

6 0.9 

i 

u 

« 

s 

.a 0.8 

o 
J3 

C 

0> 

to 

0.7 



0.6 



5 

6 
7 
I 

» 

10 
I 1 
I 2 
13 
I 4 
13 
16 
17 
11 
If 
20 
21 
22 
23 
24 
23 



0.5 



||IMhi 




. OOOE.OO 

. 443E-0I 

. tlSE-01 

. 133 

. 177 

. 221 

. 26< 

. 310 

. 334 

. 391 

. 443 

. 4J7 

. 331 

. 375 

. 620 

. 664 

. 708 

. 752 

. 7»7 

. «4I 

. «>5 

. 929 

. 974 

1. 02 

I. 06 



0.4 



0.2 



0.4 0.6 

Fraction of chord 



0.8 



1.0 



Figure 36. Swept/leaned vane design point flowfield — near-tip section. 



TE96-1049 



46 




l.i r 



1.0 



0.9 



0) 
Xi 

£ 

3 

« 0.8 



o 



§ 0.7 

to 



0.6 



0.5 \- 



0.4 



Relative Mach Number 
49.2% Span 



OOOEtOO 
0E-0I 
E-OI 




.■■■■■■B^ 





0.2 



0.4 0.6 

Fraction of chord 



0.8 



1.0 



TE96-1050 



Figure 37. Swept/leaned vane design point flowfield — midspan section. 



47 



>k 



Relative Mach Number 




9.3% Span 



10 

1 1 

1 2 
13 
14 
IS 
It 
17 
18 
19 
20 
21 

U 



. OOOE-fOO 

. 4 1 JE- 01 

. I26E- 01 

. 124 

. 165 

. 206 

. 241 

. 219 

. 330 

. 372 

.413 

. 454 

. 496 

. S37 

. 578 

. «19 

. 661 

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Figure 38. Swept/leaned vane design point flowfield — near-hub section. 



48 



Low Speed Requirements 

This inlet is required to operate at maximum takeoff flow without internal flow separation for up to 20 
degrees angle of attack (AOA) and at free stream Mach numbers ranging from to 0.25, which are typical 
of levels encountered during aircraft terminal operations. No external inlet separation requirements have 
been considered; however, it is presumed in case of engine out or shut down, the nacelle forebody cowl 
will not separate at climbing speeds with AOA below 15 degrees. No crosswind and ground operational 
requirements have been considered either. For simplicity, an axisymmetric nacelle design with zero inlet 
droop angle is assumed adequate for this application. 

High Speed Requirements 

The design cruise Mach number will equal 0.80. At the design cruise Mach number the fore and aft na- 
celle cowl contours are designed for minimal spillage and wave drag. Normally the engine nacelle is de- 
signed to have minimal total drag for a range of cruise Mach numbers, since the corresponding aircraft 
may be required to operate at different altitudes and flight Mach numbers. Generally, it is desirable to 
have a nacelle design so its overall drag remains constant or close to the design goal for flight Mach ^num- 
bers at least 5-10% above the design cruise Mach value. This upper limit of Mach number is called the 
drag divergence Mach number (M dd ). For this design, M dd is fixed at 0.86. 

Other Constraints (Geometrical) 

The nacelle aft cowl is designed to match the NASA propulsion simulator ducted prop drive rig. This 
requirement essentially sizes the overall test model dimensions, establishes the fan cowl and the core cowl 
boattail angles, and also locates the truncation point of the core cowl near the simulator metric station. 
The nacelle internal flow lines are constrained by the Allison wide chord fan design with a tip diameter of 
22.0 in. 

Since the model inlet flow and the fan duct flow are the same, the fan nozzle exit area is also sized to pass 
the maximum inlet corrected flow. Compared to the corresponding flight worthy nacelle, the fan nozzle 
is slightly larger than a scaled-up realistic fan nozzle design. The fan nozzle discharge coefficient (Cd) is 
assumed to be 0.984 (same value was used in the corresponding engine cycle) for choked flow nozzle 
conditions. No additional fan duct pressure loss has been included. 

Initially, two different inlet/nacelle designs were developed to evaluate and compare the overall nacelle 
size required to incorporate various noise suppression linings. Figure 39 compares the nacelle aeroknes 
for these configurations; however, due to program time and funding limitations a single design with a 
compact inlet and diffuser length having (L) inlet/Dff of 0.50 was selected as a baseline naceUe configu- 
ration. The selected design provides adequate surface area, or space, for advanced acoustic treatments 
both in the inlet/diffuser region and in the fan duct. The duct and cowl lengths are sufficient, when 
scaled to the reference engine size, to accommodate an advanced thrust reverser design. The geometrical 
characteristics of the baseline nacelle are presented in Figure 40. The extra-long fan duct provides enough 
space to conduct tests with alternate OGV strut designs involving a set of sweep angles and varied axial 
lengths between fan trailing edge (TE) and leading edge of the OGVs The fan and the core cowl contours 
have been designed to meet the above requirements with the external boattail angles of 10.8 and 8.8 de- 
grees, respectively, consistent with wing mounted nacelles configured for low boattail pressure drag and 
with reduced nacelle/wing interference drag. 

3.2.2 Aeroline Development 

The inlet/nacelle contours have been generated using an Allison proprietary geometry code. This en- 
abled an efficient, smooth flow path to be generated for enveloping engine hard points as weU as ^main- 
taining the desired geometrical characteristics. NASA provided the attachment hard points on the 



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simulator flow path to maintain surface continuity between nacelle aft fairings and the drive shaft. Ana- 
lytical or empirical 1-D techniques were used to provide preliminary performance projections prior to 
conducting a detailed CFD flow analysis. Figure 40 illustrates the nacelle aerolines along with important 
dimensions. 

3.2.3 CFD Analysis 

Inlet flow-field predictions using PMARC, a panel method code, were obtained to confirm the aerody- 
namic characteristics of the nacelle design. Figure 41 presents the baseline nacelle configuration analyzed, 
showing surface panels. Three flight conditions were analyzed using PMARC. These conditions are criti- 
cal to the inlet design for engine operability and maximum cruise operation, and are as follows: 

(1) Mjnf = 0.2, AOA - 20 deg, Wcorr = 102.78 lb 

(2) Mjnf = 0.8, AOA = 0.0 deg, Wcorr = 102.78 lb 

(3) Minf = 0.0. AOA = deg, Wcorr = 104.5 lb 

The analysis was conducted at several other conditions to calibrate the flow solution and the aerodynamic 
load calculation methods. Since this nacelle design will only be tested at low-speed conditions, condition 
(1) was used for the detail inlet/nacelle analyses. Typical surface flow distributions for the above condi- 
tions are enclosed in Figures 41, 42, and 43. Boundary layer analysis (Figure 44) was conducted using 
PMARC pressure distributions to provide surface skin friction Cf distribution on the inlet and nacelle to 
verify a separation-free flow. 

3.2.4 Aerodynamic Loads 

The pressure distribution obtained from the PMARC analysis at angle of attack was integrated over the 
nacelle length to obtain the resultant load and moment on the static structure due to operation at this 
condition. The results show both magnitude and point of application (fan face = 0.0) (Table II). These 
results were combined with the standard aerodynamic loads generated on the vane airfoils from deswirl- 
ing of the fan rotor exit flow to determine the structural integrity of iie static structure. 



52 




TE96-1054 



Figure 41. PMARC panels. 



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57 



4.0 STRUCTURAL DESIGN 

4.1 ROTATING COMPONENTS 

The fan rotor assembly is composed of ^^"^™ 

in g of 18 airfoils and a hub; a spinner; and a to ^ ue .^ n ^^ a pilot surfa ce and re- 

foL a bolted assembly. The blisk is P«^^y^*^S^i tJJo components through 
tained through a bolted flange arrangement. Torque is tiaD V^S^^ to[Qae sleev / to rem0 ve the 
a single sheafpin. The spinner is threaded onto me forward -P^ rf ^X5£^ tones in pre- 
need 'for attachment bolts and! *e assorted »^^^^S^a force bd^ A«£m- 
vious NASA test programs. The torque sleeve attaches to the *££*£%& fc matchin slots m ^ 
bly is by way of four cross keys ^^^J^^^^d state, was selected for the 

torque sleeve to meet the strength and life requirements of the cross keys. 
4.1.1 Stress and Deflection Analysis 

were constrained to lie within current engine experience. 
AU^siswasperformedusms™ 

symmetric analysis was performed on *e M. «* ^r^edTeSrately, with^e airfoils repre- 
tractions along the nm surface The blade stresses were ae K^ ' . , , e oliente d at 

have been determined that yield an acceptable safety factor. 

The structural audit sheets pres^ m Tab^ 

pared to material limits. Material properties contamed m dies ^J"^^^ For design as . 
prietary database and include a sufficient sample size to «^ J^*"^^^^ ( . 3 afbelow 
Lsment, the material properties used are those «««P«J^^^^Ln alloy used 

sleeve. 

AirfoU, disk, spinner, and torque ^J^£S££S 2£S?Xm^ ^ofS- 
including the appropriate ^r^^^S^^S or *e airfoil include the effects of 
tress contour plots, is presented in Appendix E. The result > presen ^ ^ 

offsetting the stacking axis axially and aI ™ I ^^^ < J^ a ^S section were required to be 
tion. Tolnsure structural integrity, ^^^^^^^i or both radial 2nd tan- 



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Figure 45. Material properties of Ti6-4 (AMS 4928). 



200.0 
TE96-1058 



300 



250 



200 




-•LOFty 
-■4CF(Kt=1) 
■A HCF (Kt = 3) 
■▼ -CF (100Q Cycles) 



S 150 

£ 
55 



100 



50 



60 80 100 120 HO ^^^o" 

Temperature, deg. F 



180 200 

TE96-1059 



Figure 46. Material properties of 17-4PH (AMS 5643): H 1100 annealed. 



62 



allowable values for the materials selected, easily satisfying the criteria. Since the blisk and torque sleeve 
form a bolted assembly, the integrity of the assembly must also be ensured. An axial stress held exists 
across the blisk to torque sleeve flange that tends to open this joint. Flange fastener sizes and assembly 
torque levels were determined based on the predicted axial stress levels across the flange to ensure sepa- 
ration will not occur. Torque transfer between the blisk and torque sleeve is accomplished through a 
dowel pin. The cross section of this pin was sized to carry the full rotor torque load at the maximum 
steady-state operating conditions in shear without help from the flange bolts. 

The burst speed corresponds to the rotational speed at which either the airfoil or disk cross section is no 
longer able to support the centrifugal loading. Standard Allison design practice requires the burst speed 
be at least 25% above the maximum steady-state operating speed of the part. For this rig, the maximum 
operating speed has been defined as 105% of the design mechanical speed, or 10,920 rpm. AUison design 
criteria are intended to ensure tensile failure will occur first in the airf oil. For gas turbine disks with cross 
sections whose thickness varies radially, failure can occur as a result of either radial or tangential over- 
load For an ideally ductile material, redistribution of the cross-sectional loading wou d occur, delaying 
failure until the full cross section reached the material ultimate strength. As a result, the primary variable 
used in assessing disk tensile failure margin is the average stress across the full d*k cross section. In cer- 
tain cases the material may not be sufficiently ductile to fully redistribute the loading, resulting in failure 
due to overstress of a local cross section. To ensure a local failure condition would not affect the burst 
margin, average tangential and radial stresses over the disk web and average radial stresses around tiie 
flange holes were also determined Referring to Table III, the limiting tensile loading in the disk for this 
design is the result of tangential stress. Little difference is observed between averaging over the fuU cross 
section or the web cross section. The predicted levels for the disk are substantially less than the 0.95 of 
tensile ultimate allowed by the criteria at 125% of maximum steady-state operating speed. The maximum 
average stress levels again occur in the airfoil hub. Referring to Table m, the design criteria require these 
average levels to be less than the tensile ultimate for the blade material at 125% of the maximum speed. 
The predicted levels satisfy these criteria. Ratioing the airfoil average stresses by the square of rotationa 
speed, burst is calculated to occur at a speed corresponding to 183% of the maximum steady-state operat- 
ing speed. 

Due to the limited running requirements for the rig, a minimum acceptable low cycle fatigue : hfc > of 1000 
type 1 cycles (idle-maximum-idle) was established. The low cycle fatigue strength for AMS 4928 and 
AMS 5659 is shown in Figures 47 and 48 as a plot of cycles to crack initiation as a function of von Mises 
equivalent stress. For the airfoil, the life critical locations are in the hub fillet and along the leading edge 
Stresses in the hub fillet were again determined through the application of a stress concentration factor of 
1 4 to the finite element results, rather than through direct calculation. Along the leading edge the effects 
of small body foreign object damage have been included through the application of a stress concentration 
factor of 2. Based on these equivalent stress levels, minimum fatigue life in excess of 1 million cycles can 
be expected. 

In addition to the stress results presented above, deflections were obtained from the finite element analy- 
sis The predicted deflections in critical areas are shown in Figure 49. At the tip, the leading edge radial 
deflection of 0.020 in. was used to set the static clearance between the outer flow-path wall and the airfoil 
to preclude rubbing over the test speed range. At the pilot surface between the blisk and torque sleeve, 
the blisk was predicted to grow an additional 0.001 in. compared to the torque sleeve, due to the differ- 
ence in elastic modulus of the two materials employed. In order to ensure accurate centering of the blisk 
on the torque sleeve at speed, this differential growth must not be allowed to open the pilot. To accom- 
plish this the mating pilot surfaces have been dimensioned to provide an interference fit at assembly. 
The predicted deflections were also used in an iterative procedure to determine the correct manufacturing 
coordinates to provide the intended aerodynamic shape at the design speed. The coordmates for the air- 
foil at static, 85% N d , and N d are tabulated in Appendix B. 



63 




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8 100 




10" 10' 

Cycles to Crack Initiation 



10° 10* 

TE96-1061 



Figure 48. Low cycle fatigue strength of AMS 5659 (17-4PH) at 70°F. 
4.1.2 Vibration Analysis 

Vibration analysis of the integral bladed disk was carried out to define potential areas of vibratory re- 
sponse and to ensure adequate high cycle fatigue strength was available to allow operation over the entire 
design speed range. Specific consideration was given to avoidance of flutter over the rig operational en- 
velope, placement of potential resonant conditions in speed ranges away from where substantial test time 
was to be accumulated, and satisfaction of minimum fatigue strength requirements over the entire bladed 
disk 

Natural frequencies of the bladed disk system were obtained from finite element analysis at a series of 
rotational speeds. The finite element model consisted of a single airfoil supported on a pie-shaped sector 
of the disk. The periodic structure of the system was retained through application of cyclic symmetry 
boundary conditions along the edges of the disk sector. The airfoil was represented by a mesh of 8-node 
meanline shell elements, while the disk was modeled with 20-node solid elements. For completeness, 
comparisons of natural frequency and mode shape were made between the full bladed disk model and a 
cantilevered airfoil model. This comparison indicated insignificant levels of disk participation in the vi- 
bratory modes. 

The results of the natural frequency analysis, in the form of a Campbell diagram, are presented in Figure 
50. Plots of the deflected mode shape and resulting vibratory stress distribution are found in Appendix F. 
The diagonal engine order lines represent the locus of excitation frequencies produced by flow asymme- 
try with wavelength corresponding to the order number. Low order excitation (i.e., 2, 3, or 4EO) is typi- 
cally the result of inflow total pressure or temperature distortion. Allison development experience indi- 
cates the coincidence of the fundamental bending (IB) and torsion (IT) natural frequencies with second 
and third-engine order should be avoided in speed ranges where significant operational time will be 



65 



0.02012 f 001430 a 0.00101 




0.00335 



u 

TE96-1062 



Figure 49. Predicted rotor radial de flections. 



66 








1 



A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design 



4 5 6 7 8 
Rotor Speed, krpm 



10 11 



TE96-1063 



Figure 50. Campbell diagram of blade. 



67 



SS2 ■£££££! (5 2» imfLST' deSig " r *■ SpeedS «»»««* to Ml power 

provide a minimum 15% speed margin relative tn ? « jSr? » £ eS Were ad us *ed to 

cem was excitation of higher ordTmodes bv le vt^lS * J** 6 "^ Speeds - « secondar y «»- 
this stage, resulting in potential excrTation at JS^S £?"* ^ P ? ^ "' Tbm are 42 vanes * 
15% speed marginletween all n^ZZf^SEol^Zt of ^ "" P ^ 6 10 ^^ 3 
diction of the resonant response of a mode to eTcLtion^h as ^ ^ ^"T 7 "^ AcCurate pre " 
tificarion of specific modes whose resorumt 12!" n k adUCVed yet ' fcus P r «^ding iden- 

was thus to minimize the number of mo£s exnenVn acceptably large. The design strategy 

mterest. As finally ^ep^^sp^Ue^;^! wTeeT^ ^ ** * We > eed8 S 
except the 21st mode. Due to the relatively e^Z soL n^?!^ "P™ betWeen 42EO *"* ^ m ° d « 
tor leading edge, a weak excitation Ao^^^tl^^T *f T '"^ ^ md ** Sta " 
^ofsimnarcomponentswimsimnarroto^-^ 

onset. An analytical method, wSS^^ a ^^3f ,P ^ d * ^ ** ^^ ° f fl " tter 
deflection pattern, is available and h£ So^en hiSvlSaT SET* T^u ^ ' ^^ m ° dal 
cal formulation, it is only applicable in SpersoSow, ^ 7f ' ^ t0 ** methods ma th e mati- 

relative Mach numbers too low for aS2Srf t£ I*, P ? ^ ** Sp6ed r6SultS m M « 

an empirical correlation has be^d^^^^nlZ T ° " U8 T lt ** anaJ > rtical method ' 
the product of chord»fr eq uency/(2- ^ ° r reduCed ^^^ defir * d as 

established at 0.2 for meLda^LaTbldrng mod^ So 6 Sffi ^j 1 "™ vaI -s) have been 
motion. For the current design, the calculated rTdurp^l ?? m0de Wlth s^^* torsional 

0-72. These satisfy the criteril CalCUiated reduCed *«F«naes of the relevant modes are 0.29 and 

reversed bending that can be Sosedon a mate^T V"*? " ** Vibrat ° ry StreSS level m *% 
endurance strenfth is reduced™ hen a meanTtreS fi^H ^^f ^ ^ *«*" faiIures - ^ 
ing zero as the n£an stress apprTacnes ^eTenS uM™?/^' "^ ^ enduranCe 8te «* a PP roa <*- 
graphically in the Goodman dlgra^ l^ord"! tTet J^S^l ^^ * *P kaJ * P~ Sented 
in fatigue, Allison requires the rninimurn ZhZT ,, a f asonabe vibratory response will not result 
the airfoil, after accoItLg For SS^ZSSl^ T" * " ^ 15 ** ^ aU l0Cati ° ns on 
fatigue data for notched s^ne^ftSSSS^ ? w" ^f ^ * aSS6SSing ^ Criterion ' 

hub and edge regions le^v^t^o^^SlJSf T*'??' ons ' k t ' of " and 2.0 are used in the 
fatigue dataware based Jj^£££££% J*** £?!? ^ <*** da ™ge- In other regions, 
^wablestressof^immeleadmg^^^^^ 



4.2 STATIC COMPONENTS 



of spool segments forming the internal flow nJ^!^ f 0lk m each ^S" 1 **- ^d a series 

effects of va^e geometry, Ltp^te^^^^^ 

carrying member. The vane segments are tied fn ^Tcf , I !. S * e Vanes to become a Ioad 

shear pin. The shear pin provtdS me S£t L^jSf ^^ S t PP ° rt by *« bolts ^ d a °250-in. 
vane segment ag^V»t.tic^^rdto^SSS^; , ;' hiIe ** "^ faSt6nerS ^ ^ 

andva.eassembhesarecLLctedSS^^^^^ 



68 




!S>j 'ssajjg AjojejqjA 



69 



soecification to provide the required strength and rigidity. The flow-path spool pieces are constructed of 
SSSdk^Ss 4127 (6061) in the T6 condition, to minirnize . >verhung weight Weight reduction 
w^Tp^rity To n^Te Tig deflection at the blade track and t > facilitate handling durmgassemby. 
StionaSterZ^th pieces have been designed to adapt the rig to an existing *^*«£ var - 
aMe area nozzle allowing stage performance measurements to be acquired. To deal with the additional 
SSSSJLL irirtion of these pieces, provisions for external support have been pro- 

vided. 

4.2.1 Stress and Deflection Analysis 

Structural analysis was carried out for each of the vane configurations at two loading conditions. The first 
SI S£ Resents standard rig operation and consists * the nacelle weight and aerodyna^c 
£S derated on the vanes as they deswirl the rotor discharge flow. In the second condition addi- 
ti^aTLrSy^aScToads arfapplied as a result of operating the nacelle at an angle of attack to the wmd 
^el flow SrTctural assessment criteria employed to evaluate the integrity of the static components 
Sowed sTanZd Allison practice for nonflight applications. Specific consideration was given to tensde 
Jm^tSktfelding, creep, low cycle fatigue, and deflection resulting from nonaxisymrnetnc loading. 
S^toftetoSS Me, Learc^ nature of the rig, no provisions for containment m the event of an airf oil 
S were^cluded in the design of the nacelle. For this reason, human proximity to the rig during op- 
eration should be avoided. 

All structural analysis was performed using the finite element method. A model composed of a 1/42 
icto- *e entJstatic stature, corresponding to a single v-e passage, was generated f^eachof £e 
fir vane configurations for analysis in the Allison proprietary FEM procedure, STRATA. The vane inner 
cTdTas deSed using 20-node solid elements. Beam elements were employed to represent the m- 
n«ba^d attaint bolts, mner band shear pin, and the attachment bolts in the outer flanges. The rest of 
Z s^uctoe ^modeled with 8-node meanline shell elements. The static structure att achment to 
^ouS was through two spring elements at the pilot surfaces representing the rig static balance stiffness. 

As previously mentioned, structural analysis of the rig was based on ^^^^^^ 
nf the rie at an angle to the tunnel flow produces a nonaxisymmeinc loading on the nacelle. «^ omc 
Toadmg of thTsSr model was used to account for this asymmetry. As a result of the asymmetric .load 
rSaCSsand deflection patterns are also asymmetric. For nonsymmetnc loadrng conditions, 
Sc^Tcn^eria are a^essed at the worst location in the assembly. At the edges of the modeled sector, 
^^^b^^onditions consistent with a split hoop are applied along the faces of the inner 
Id ouSTids A secondary result of applying cyclic symmetry over a single vane passage W*. is 
that the model represents a structure with one bolt and one shear pin for each airfou. This modeling inac 
^racv w^ notXt the stress and deflection field away from the attachment points and was used to re- 
d^e compute resource requirements. To assess the stresses in t,e shear pin and attachment bolts, it was 
fSSd^SSS additional constraints results in an equa' increase in load * *7~S ™ m " 
Srr-This^oduces a factor of six increase in the section stresses in the shear pm and a factor of two in- 
^^bohSe^ This approach is not entirely accurate, but the resulting stresses are so low that 
a^L^-P^Si was deemed unnecessary. The bolted flanges on the outer duct P^ .way 
torn 11 vSe w2e not represented in detail in the finite element model. Bending stresses in die flanges 
were dSermSed by hand calculation. A conservative approach was taken, requiring a single flange seg- 
meS betw^rwo bolts to carry the entire nacelle bending moment due to angle of attack operation. 

The structural audit sheet (Table V) summarizes the results of th e analysis relative to the design criteria. 
S p^S t^ctrTlconce^ 

assessed using equivalent stress as defined by the Von Mises en ena. Referring to Table V the peak 
eauTvalent Ss occurs in the baseline vane hub trailing edge filet when this vane is installed in the aft 
Sn i^re £ ™s stress is 41% of the material yield, whi, h satisfies the Allison cntena. As axial 
P wip S SSuced, the peak stress levels decrease. This is a result of changes in the load transfer 
m^ch^sm between the configurations. For the baseline vane, which has a radxal stack axis, the nacelle 



70 



loading is reacted out by the vane in pure bending about an axis normal to the airfoil plan view. This re- 
sults in the majority of the load being transferred along the leading and trailing edge. As sweep is intro- 
duced, a portion of the nacelle load is transferred as tension parallel to the stack axis, similar to diagonal 
members in a truss. Since the section structural efficiency in tensile loading is greater than for bending, 
the resulting peak stress is reduced. 

Table VI shows the circumferential variation in peak stress due to the load asymmetry for each of the 
vane configurations. Also shown in the table is the maximum stress due to the normal aerodynamic de- 
swirl loads. Complete results of the stress analysis, in the form of isostress contour plots, is presented in 
Appendix G. Referring again to the audit sheet, the maximum stress in any of the flanges is found to be 
7.5 ksi. These flanges are retained with 34 fasteners with 0.190-in. diameter. Standard torque levels for 
these fasteners will be sufficient to prevent opening of the flanges. The stress levels shown for the fasten- 
ers on the vane inner band reflect the Allison design practice of preloading fasteners at bolted joints to 
80% of the material yield. In this application, the fastener stress is composed of 57 ksi due to preload and 
a 23 ksi bending stress from the vane loading. As in the rotating components, the design goal for low cy- 
cle fatigue life was 1000 type 1 cycles (minimum). Crack initiation is governed by local stress peaks; thus, 
the vane hub trailing edge fillet stress of 56 ksi will set the life potential for the static structure. The vanes 
are constructed from wrought 17-4PH stainless steel. Since the limiting stress occurs along an edge, a 
theoretical stress concentration of 3 is applied for life assessment to account for possible small object for- 
eign object damage in this area. Based on these assumptions, the predicted low cycle fatigue life is 66,000 
cycles. 

In addition to the stress field induced in the vane and nacelle structure, operation at angle of attack will 
produce a deflection of the casing relative to the blade tip. The design is intended to have a uniform 
running clearance of 0.020 in. at the design rotational speed. The casing deflection at the blade track due 
to the nacelle loads is tabulated for the various vane configurations in Table VII. A maximum radial de- 
flection of 0.006 in. is predicted and will occur in the swept and leaned configuration. Complete plotted 
results of the deflection analysis for both load conditions are presented in Appendix H. 

4.2.2 Vibration Analysis 

Vibration analysis of the static structure was carried out to define potential areas of vibratory response 
and ensure adequate high cycle fatigue strength was available to allow operation over the entire design 
speed range. Specific consideration was given to avoidance of flutter over the rig operational envelope, 
placement of potential resonant conditions in speed ranges away from critical test speeds, and satisfaction 
of minimum fatigue strength requirements over the entire structure. 

Natural frequencies of the static structure assembly were obtained from finite element analysis. A finite 
element model of a 1/42 sector of the structure was generated for each of the four vane configurations 
and for both the flight and performance measurement ducting arrangements. Above the third natural 
mode, deflections tend to isolate in the vane assembly. To reduce the computational requirements, a re- 
duced order finite element model representing the airfoil and vane outer and inner band was constructed 
to obtain these higher modes. Comparisons of the full system and reduced order models for a limited 
number of modes substantiated the accuracy of this approach. In the performance configuration, weight 
isolation for the inlet bellmouth and variable area nozzle will be provided. Based on the methods under 
consideration for providing this weight isolation, it was assumed they would not contribute to the system 
stiffness. The connection between Allison's static structure and the rig static force balance was simulated 
by springs at the pilot locations, with spring rates obtained from Boeing design documentation provided 
by NASA. Natural frequencies and mode shapes were calculated using the Allison finite element code, 
STRATA. Calculations of system response to unbalance were carried out using the Allison forced re- 
sponse code MODLRESP, running as a post-processor to STRATA. 



71 





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Table VI. ... 

xt*<:a ^.H fan rig n e~"- — ^ *™ summaries (All values Vo n Mises e quivalent stresses fksil) . 



Vane 



^ • *: TjpLE TipTE 

Description bi MO 

Maximum stress due to vane loads 9.1 44-u 

Maximum stress due to vane + AOA loads + nacelle weight 



Vane 1, 90 deg 
Vane 6, -135 deg 
Vane 11, -180 deg 
Vane 17, -225 deg 
Vane 22, 270 deg 
Vane 27, -315 deg 
Vane 32, -0 deg 
Vane 38, -45 deg 



12.5 
10.6 
8.4 
7.9 
8.0 
7.9 
9.1 
11.9 



27.2 
28.8 
35.1 
44.6 
48.8 
46.6 
39.7 
30.9 



HubLE 
36.3 

53.5 
50.4 
40.7 
28.1 
23.2 
25.9 
35.2 
48.3 



HubTE 

48.3 

29.9 

33.5 
41.9 
51.1 
54.2 
51.8 
44.4 
33.8 



Vane 



Description TipLE 

Maximum stress due to vane loads 8.8 
Max stress due to vane + AOA loads + nacelle weight 

Vane 1, 90 deg JJ.4 

Vane 6, -135 deg ai1 

Vane 11, -180 deg 9 _ 2 

Vane 17, -225 deg "_ 7 

Vane 22, 270 deg 7 _ l 

Vane 27, -315 deg J.l 
Vane 32,-0 deg 

Vane 38, -45 deg llA 



TipTE 
43.2 

25.8 
26.6 
33.1 
44.0 
49.2 
47.3 
40.0 
30.3 



HubLE 

32.7 

48.6 
46.7 
37.9 
24.7 
18.9 
21.3 
30.5 
43.1 



HubTE 
48.6 

29.2 
31.8 
40.3 
52.0 
56.7 
53.9 
45.2 
33.8 



Vane 



^ • ^ TjpLE TipTE 

Description -tea 62 

Maximum stress due to vane loads 35.8 °- 

Maximum stress due to vane + AOA loads + nacelle weigh 



Vane 1, 90 deg 
Vane 6, -135 deg 
Vane 11, -180 deg 
Vane 17, -225 deg 
Vane 22, 270 deg 
Vane 27, -315 deg 
Vane 32, -0 deg 
Vane 38, -45 deg 



43.0 
38.5 
32.7 
29.2 
29.5 
32.0 
36.7 
42.5 



14.1 
12.6 
8.3 
3.6 
3.6 
3.7 
7.1 
12.3 



HubLE 
24.1 

34.3 
31.2 
25.0 
18.2 
16.4 
19.2 
25.3 
32.5 



Bolt Shear pin 
16.0 15.0 



13.8 
13.6 
14.4 
16.0 
16.6 
16.4 
15.4 
14.4 



HubTE 
14.2 

8.2 
9.2 
15.3 
24.2 
26.2 
22.8 
14.6 
9.2 



Vane 



r. • ^ TjpLE TipTE 

Description "^' 1T7 

Maximum stress due to vane loads 32.8 .,7 
Maximum stress due to vane + AOA loads + nacelle weight 

Vane 1,90 deg 24.0 ■ 

Vane 6, -135 deg ^U.l • 

Vane 11, -180 deg 36.1 »•* 

Vane 17, -225 deg ^.5 16.3 

Vane 22, 270 deg ^4.5 17.8 

Vane 27, -315 deg 30 " 4 133 



HubLE 

7.4 

7.2 

3.0 

5.4 

14.4 

16.3 

12.0 



HubTE 
46.5 

33.1 
37.0 
43.8 
49.3 
49.8 
47.4 



18.6 
16.2 
15.0 
16.8 
17.4 
15.6 
15.0 
18.0 



Bolt Shear pin 
17.8 14.4 



16.6 
16.4 
16.8 
18.0 
18.4 
18.0 
17.4 
16.8 



19.2 
16.8 
13.8 
15.6 
17.4 
16.2 
16.2 
19.2 



Bolt Shear pin 
11.8 15.6 



7.6 

8.0 

9.6 

13.2 

15.4 

14.2 

11.0 

8.8 



18.6 

17.4 
16.8 
18.0 
18.6 
16.2 
14.4 
17.4 



Bolt Shear pin 
14.4 16.2 



3.4 
4.6 
11.2 
20.4 
22.8 
18.8 



18.0 
16.8 
17.4 
6.6 
18.0 
15.6 



73 



Description 
Vane 32, -0 deg 
Vane 38, -45 deg 



Table VI front) 

IiE-LE TipTE H ubLE Hub TE 

260 5.6 3.6 42.3 

22-3 2.1 5.5 35.1 



Bolt Shear pin 

10.4 16.2 

3.0 18.6 



Notes: 



All values are Von Mises equivalent stresses (ksi) 

degrees is top dead center, with angle increasing counterc ockwise (aft looking forward) 



Table VEL 
NASA scaled fan rig nacelle blade track dPfWHnn c ,„^ arv ( defWnW in WW) 



Description 

Maximum deflection due 
Vane 1, 90 deg 
Vane 6, -135 deg 
Vane 11, -180 deg 
Vane 17, -225 deg 
Vane 22, 270 deg 
Vane 27, -315 deg 
Vane 32, -0 deg 
Vane 38,-45 deg 



Radial 

to vane + AOA 

4.260e-03 

3.350e-03 

6.120e-04 

-2.910e-03 

-4.190e-03 

-3.270e-03 

-6.280e-04 

2.880e-03 



Baseline vanp 
Tangential 
loads + weight 
-7.030e-02 
-7.310e-02 
-7.470e-02 
-7.360e-02 
-7.080e-02 
-6.880e-02 
-6.780e-02 
-6.850e-02 



Axial 

-5.620e-0:i 
-6.820e-03 
-l.OlOe-02: 
-1.430e-02 
-1.580e-O2 
-1.460e-02 
-1.140e-02 
-7.190e-03 



Radial 

4.940e-03 
4.260e-03 
1.220e-03 
-3.410e-03 
-5.290e-03 
-4.100e-03 
-8.100e-04 
3.230e-03 



Aft Vane 
Tangential 

-6.940e-02 
-7.290e-02 
-7.500e-02 
-7.400e-02 
-7.050e-02 
-6.770e-02 
-6.630e-02 
-6.720e-02 



Axial 

-5.660e-03 
-6.570e-03 
-9.510e-03 
-1.350e-02 
-1.500e-02 
-1.410e-02 
-1.110e-02 
-7.240e-03 



Description 

Maximum deflection due 
Vane 1, 90 deg 
Vane 6, -135 deg 
Vane 11, -180 deg 
Vane 17, -225 deg 
Vane 22, 270 deg 
Vane 27, -315 deg 
Vane 32, -0 deg 
Vane 38, -45 deg 



Radial 

to vane + AOA 

3.420e-03 

3.170e-03 

6.240e-04 

-2.960e-03 

-4.220e-03 

-3.010e-03 

-2.790e-04 

2.820e-03 



Swept vanp 
Tangential 
loads + weight 
-5.820e-02 
-6.100e-02 
-6.250e-02 
-6.140e-02 
-5.850e-02 
-5.690e-02 
-5.620e-02 
-5.690e-02 



Axial 

5.000e-03 
3.340e-03 
-7.120e-04 
-5.620e-03 
-7.200e-03 
-5.540e-03 
-1.530e-03 
3.390e-03 



Swept and lpaned vanp 
Radial Tangential Axial 



5.740e-03 
4.850e-03 
1.295e-03 
-3.890e-03 
-5.990e-03 
-4.720e-03 
-1.010e-03 
3.730e-03 



-1.710e-02 
-2.120e-02 
-2.371e-02 
-2.240e-02 
-1.860e-02 
-1.610e-02 
-1.460e-02 
-1.500e-02 



1.340e-03 
-1.100e-03 
-6.668e-03 
-1.320e-02 
-1.510e-02 
-1.270e-02 
-7.160e-03 
-6.390e-04 



r^e e sented tS fo°r f th C f*™* ^"T^ ^^ ° f *" ** S y Stem ' m ** lom of » CampbeU diagram are 

It was not possible to adjust the frequencies of these modes sufficient!) to mov fte reson^ con^tioT' 
ou side the test speeds of the rig. Since both modes produced a result- nt radTa! defied W unTblade 
££Z TT^rT I ^f^ C ° Uld feSUlt " COntact between * ' "** tips andWasl To de- 

oXpt dt e otL° e ^ th , « ^ 3 ' ° rCed reSP ° nSe "** Wa£ COnduc * d --S aShZ^e 
toad applied I in phase at the static structure support pilot surfaces. A damping of 6.3% (loe decrement 

was assumed a conservative assumption based on Allison experience. The resulting bladf n-STflec 

tions for the pitch mode, which is the most sensitive to excitation, is pr, :sented in ££« 56 



74 




1 2 



3 4 5 6 7 
Rotor Speed, krpm 



8 



10 11 



A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design TE96-1065 

Figure 52. Campbell diagram for baseline vane in acoustic testing setup — assembly modes. 



75 



200 



12 3 

A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design 



4 5 6 7 
Rotor Speed, krpm 



1 EO 




10 11 



TE96-1066 



Figure 53. Campbell diagram for aft vane in acoustic testii ig setup — assembly modes. 



76 




1 2 



A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design 



34567 8 910 11 
Rotor Speed, krpm 



TE96-1067 
Figure 54. Campbell diagram for swept vane in acoustic testing setup — assembly modes. 



77 



200 



180 - 



160 



140 - 



ts) 

o 

c 
o 

cr 
o 



120 



100 



80 - 



60 



40 



20 











Idle 

Approach 

Full Power Takeoff 

Design 



4 5 6 7 
Rotor Speed, krpm 



1 EO 




10 11 



TE96-1068 



Figure 55. Campbell diagram for swept and leaned vane in acoust c testing setup — assembly mode. 



78 



10" 



c 
<u 

E 
o 

C3 

a. 

sio- 

o 



f- 

CQ 

5 



10" 




Baseline Vane 



Aft Vane 
Swept Vane 

Swept & Leaned 
Vane 



Max. Static 
Deflection 



10 



Fan Unbalance, lb-in 



TE96-1069 



Figure 56. Blade track radial deflection versus fan unbalance — pitch mode of acoustic testing setup. 



79 



Accounting for the 0.005 in. (worst case) of static deflection occuring during angle of attack operation, a 
minimum unbalance of 4 in. -lb would be required to produce a rubbing condition for this mode. This 
level is two orders of magnitude larger than Allison balance requirements for hardware of this size. 
When configured in the performance mode, only one mode, labeled fore and aft in the Campbell diagram, 
coincides with 1EO within the steady speed range, Figures 57 through 60. A response calculation showed 
a residual unbalance greater than 10 in. -lb would be required to produce rubbing in this instance (Figure 
61). The Campbell diagrams for the higher frequency modes, invoh ing motion of only the vanes, are pre- 
sented in Figures 62, 63, and 64 and correspond to the four test configurations. Since these modes involve 
vibration of only the vane segments, the results are independent of the nacelle configuration and do not 
change when the radially stacked airfoil is moved into the aft position. Since the rotor contains 18 blades, 
the primary concern for resonant vibration is the placement of the 18EO coincidences with the natural 
modes. Allison experience with fixed geometry vanes indicates resonant excitation of the fundamental 
bending, or IB, mode should be avoided in the steady-state speed range. For all configurations, 1B-18EO 
resonance occurs well below the test speed range. This resonance should impose no restrictions on the 
test program. Three other modes are predicted to encounter resonant excitation within the steady-state 
speed range. The fundamental torsion (IT) and second bending (2B) modes exhibit a coincidence with 
18EO at part speed conditions. For both of these modes at least a 15% speed margin exists between the 
resonant speed and the speeds at which the primary acoustic data will be acquired. Should an unexpect- 
edly high response be observed in either of these modes, a modification to the test matrix to avoid the 
resonance can be implemented without compromising the test objectives. The second torsion (2T) mode 
of the vanes is also susceptible to an 18EO resonance. This resonance is predicted to occur approximately 
5% below the design speed for the two swept configurations and at die design speed for the baseline con- 
figuration. Accurate prediction of aerodynamically induced resonant vibration levels remains beyond the 
state of the art. Review of recent Allison vane design experience re\ eals a number of successful core 
compressor stages have similar occurrences. In these stages, the measured response of the second torsion 
mode has been uniformly low. Since the present rig employs a mucn larger spacing between the rotor 
and stator than possible in a core stage, no unacceptable vibratory response of the 2T mode is expected 
and no attempt was made to change its natural frequency so as to avoid the 18EO resonance. Plots of the 
deflected mode shapes and resulting vibratory stress distributions are provided in Appendix I for the 
system modes and Appendix J for the vane modes. 

While a relatively rare occurrence for a vane, avoidance of flutter throughout the operational range must 
be ensured. Allison has developed an empirical criterion for flutter avoidance based on reduced fre- 
quency as described in the rotating components section. Empirical "imits for minimum acceptable values 
have been established at 0.2 for the fundamental bending mode and 0.6 for the fundamental torsion 
mode. The calculated reduced frequencies for the relevant modes for each of the vane configurations is 
presented in Table VIII. All configurations satisfy the requirements 



80 




1 

A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design 



Rotor Speed, krpm 



TE96-1070 



Figure 57. CampbeU diagram for baseline vane in performance calibration setup - assembly mode. 



81 



1 

A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design 



Rotor Speed, krpm 



1 EO 




10 11 



TE96-1071 
Figure 58. Campbell diagram for aft vane in performance calibration setup - assembly mode. 



82 



A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design 



3 4 5 6 7 
Rotor Speed, krpm 



1 EO 




11 



TE96-1072 



Figure 59. Campbell diagram for swept vane in performance calibration setup — assembly mode. 



83 



x*+v 


A 


B 


C 


D 




220 


Pitch 










200 












180 










1 EO 


160 












N 140 


- 




/ 






o 


Fore & Aft 




/ 






Frequer 

8 § 






/ 










/ 






80 












60 


/ 










40 


Torsion / 










20 


- Pitch / 








n 


/ . i . i . i . i 


i 


i.i.i. 


i . i 



1, 

A: Idle' 

B: Approach 

C: Full Power Takeoff 

D: Design 



2 3 4 5 6 7 
Rotor Speed, krpm 



8 9 10 11 



TE96-1073 



Figure 60. Campbell diagram for swept and leaned vane in performance calibration setup — assembly 

mode. 



84 



1(T 



10" 




c 

CO 

E 

O 

C3 



O 



H 

.« 
CO 



10" 



10 



10" 






Fan Unbalance, lb-in 



Max. Static 
Deflection 




Aft Vane 

Swept & Leaned 

Vane 

Swept Vane 



Baseline Vane 



TE96-1074 



Figure 61. Blade track radial deflection versus fan unbalance — fore and aft mode of performance cali- 
bration setup. 



85 



4 - 



3 - 



o 

c 

4) 

P 2 



1 - 








1 

A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design 



4 5 6 7 
Rotor Speed, krpm 



18 Blades 



1 EO 



10 11 



TE96-1075 



Figure 62. Campbell diagram for baseline and aft vanes — airfoil modes. 



86 



12 3 

A: Idle 

B: Approach 

C: Full Power Takeoff 

D: Design 



4 5 6 7 
Rotor Speed, krpm 



18 Blades 




1 EO 



8 9 10 11 



TE96-1076 



Figure 63. Campbell diagram for swept vane — airfoil modes. 



87 




Idle 

Approach 

Full Power Takeoff 

Design 



4 5 6 7 8 
Rotor Speed, krpm 



18 Blades 



1 EO 



10 11 



TE96-1077 
Figure 64. Campbell diagram for swept and leaned vanes — airfoil modes. 



discussed in the rotating components secfon. ^^vdX^e MuTtos requirement must be sat- 
vane configurations satisfy the criteria. 



Tnnfi juration 

Baseline 
Aft vane 
Swept vane 
Swept and leaned 



Table VIII. 
Flutter par?™pter vane rnnfiflurations. 



IB -Hz Uzlfa 75% chord Vjdodty^sec IBJreduced) JT (reduced) 



820 
820 
619 
636 



1320 
1320 
1362 
1428 



1.810 
1.810 
1.500 
1.500 



761 
761 

774 
774 



0.51 
0.51 
0.31 
0.32 
>0.2 required 



0.82 
0.82 
0.69 
0.72 
>0.6 required 



10 20 30 40 50 CO 70 80 90 

Steady Stress (ksi) 



Material: 17-4PH 
120° F 

1 0x1 6 cycles 
-3c properties 




100 110 120 130 140 150 
TE96-1078 



Figure 65. Goodman diagram for baseline vane. 



89 




Material: 17-4PH 
120° F 

1 0x 10 6 cycles 
-3o properties 



T~ •—*, r*-, ,_ 






10 2b 30 4b 50 'e'o 70 80 ' 90 ' 




Steady Stress (ksi) 
Figure 66. Goodman diagram for aft vane. 



100 110 120 130 140 ' 150 
TE96-1079 




Material: 17-4PH 
120° F 

10x10 6 cycles 
-3c properties 






10 203040 lb~^0 




7 80 90 1(0 110 120 130 140^50 
Steady Stress (ksi) 50 



TE96-1080 



Figure 67. Goodman diagram for swej .t 



vane. 



90 



$? 



Material: 17-4PH 
120°F 

10x1 s cycles 
-3a properties 




10 2 ° 30 40 50 6b '^b~~lb~ 

Steady Stre 

Figure 68. Goodman diagram for 



S-eady Stress^ 1 °° "° 12 ° 13 ° ^ & 



TE96-1081 



swept and leaned vane. 



91 



APPENDIX A 



THE BASELINE LOW NOISE FAN: 

AERODYNAMIC DESIGN POINT 

BLADE AND VANE ELEMENT PERFORMANCE AND GEOMETRY OUTPUT 



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117 



APPENDIX B 



LOW NOISE FAN CONFIGURATION NO. 2: 

AERODYNAMIC DESIGN POINT 

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181 



APPENDIX E 
STRUCTURAL ANALYSIS RESULTS ROTATING COMPONENTS 




_x 



TITL E NASA ».. F.N DEFAULT BC" S HOT- TO- COL. , M. C, LNF. FNL - >....« 

GEOMETRY PLOT ,,.,, 94/154 

SCALE - 1.0000 PLOT TIME AND DATE - 11=30.25 94/154 



e Sor face 



185 




— X 



• » • 


LEGEND • 




KS 1 


A 


32.00 


B 


27. 00 


C 


22. 00 


D 


17.00 


E 


1 2. 00 


F 


7. 00 


G 


2. 00 


MAX 


32. 75 


MIN 


. 54 



TITLE NASA 22 i. FAN DEFAULT BC" S HOT- TO- COLD [J..C) LNF FNL 
CONTOUR PLOT OF VON MISES UNIAXIAL EQUIVALENT STRESS 
SCALE- I. 0000 PLOT TIME AND DATE . 11:30:13 94/154 



Pt«ss*r< Sarface 



186 




_x 



» • • 


LEGEND •• 




KSI 


A 


11.00 


B 


9. 00 


C 


7. 00 


D 


5. 00 


E 


3. 00 


F 


1 . 00 


G 


-1.00 


H 


-3.00 


MAX 


11.30 


MIN 


-4.91 



TITLE NASA 2 2i. FAN DEFAULT BC S HOT- TO- COLD . M. C, LNF. FNL 
CONTOUR PLOT OF SIGMA X COMPONENT STRESS 
SCALE = 1.0000 PLOT TIME AND DATE - 11:30:39 94/154 



•DENOTES HIDDEN 



Fmuit S»r f «e« 



187 




• • * 


LEGEND • 




KS 1 


A 


3. 00 


B 


2. 00 


C 


1 . 00 





. 00 


E 


-1.00 


F 


-2.00 


G 


-3.00 


H 


-4.00 


I 


-5.00 


J 


- 6. 00 


K 


-7.00 


MAX 


3.15 


MIN 


-7.01 



'DENOTES HIDDEN 



TITLE NASA 22 i. FAN DEFAULT BC'S HOT- TO- COLD I Mi CI LNF FNL P, . . 

CONTOUR PLOT OF STOMA Y COMPONENT STRESS ' """" *"""< 

SCALE - I. 0000 PLOT TIME AND DATE . 11:30:43 »4/lJ 4 



188 




__x 



• * • 


LEGEND • 




KS I 


A 


3 1.00 


B 


24.00 


C 


17. 00 


D 


10.00 


E 


3.00 


F 


-4.00 


G 


-11.00 


H 


-18.00 


MAX 


3 1.30 


MIN 


-20.05 



TITLE NASA 22 i. FAN DEFAULT BC" S HOT- TO- COLD I Mi C] LNF. FNL 
CONTOUR PLOT OF S IGMA Z COMPONENT STRESS 
SCALE = 1.0000 PLOT TIME AND DATE « 11:30:47 94/1S4 



Piiii>[< Sir f »ce 



189 




__x 



TITLE NASA 22 i. FAN DEFAULT IC'S HOT- TO- COLD [MCI LNF.FNL 
CONTOUR PLOT OF MAXIMUM PRINCIPAL STRESS 
SCALE = 1.0000 PLOT TIME AND DATE - 11:30:50 94/154 



• * • 


LEGEND ••• 




KS I 


A 


33. 00 


B 


28.00 


C 


23. 00 


D 


1 ». 00 


E 


1 3. 00 


F 


>. 00 


G 


3. 00 


MAX 


33.29 


MIN 


- . 03 


Prenire 


Sir face 




I 



190 




» * * 


LEGEND * 




KS 1 


A 


11.00 


B 


9. 00 


C 


7. 00 


D 


5 . 00 


E 


3. 00 


F 


1 . 00 


G 


-1.00 


MAX 


11.04 


MIN 


- 2. 23 



TITLE NASA 22 i. FAN DEFAULT BC 1 S HOT- TO- COLD ( M.CI LNF. FNL 
CONTOUR PLOT OF SECOND PRINCIPAL STRESS 
SCALE * 1.0000 PLOT TIME AND DATE « 11:30:33 94/154 



Ptcitirc Sir! ice 



191 




_x 



• • • 


LEGEND • 




KS I 


A 


. 00 


B 


-4.00 


C 


- ». 00 


D 


-12.00 


E 


-16.00 


F 


-20.00 


G 


-24.00 


MAX 


. 01 


MIN 


- 27. 01 



TITLE NASA 22 i a FAN DEFAULT BC'S HOT- TO- COLD I M» C] LNF.FNL 
CONTOUR PLOT OF MINIMUM PRINCIPAL STRESS 
SCALE - 1.0000 PLOT TIME AND DATE - 1:30:56 94/154 



Preinf • Svrface 



192 




TITLE K. SA ».. - — BC-S HO T -TO-CO LD .-C, — - — » 



Sir face 



193 



X 




CHoL L r E R r" L A oT 2i • FAN ^^ BCS H0TT — '>- ™.»t. ■ -«. ... ..,..«. 

SCALE- 1.0000 PLOT TIME AND DATE . I1:30:S7 94/154 



194 



x_ 




• w • 


LEGEND •' 




KS 1 


A 


32.00 


B 


27 . 00 


C 


22.00 


D 


17.00 


E 


12.00 


F 


7.00 


G 


2. 00 


MAX 


32.75 


MIN 


. 54 



TITLE NASA 2*1. FAN DEFAULT .C S HOT- TO- COLD IM.C1 LNF. FNL - S.c, io. Sorf.ce 
CONTOUR PLOT OF VON MISES UNIAXIAL EQUIVALENT STRESS 
SCALE = I. 0000 PLOT TIME AND DATE - 11:31:03 94/154 



195 




* • • 


LEGEND • 




KS 1 


A 


11.00 


B 


9. 00 


C 


7. 00 


D 


5.00 


E 


3. 00 


F 


1 . 00 


G 


-1.00 


H 


-3.00 


MAX 


11.30 


•MIN 


-4.91 



TITLE NASA 22 ii FAN DEFAULT EC'S HOT- TO- COLD I 
CONTOUR PLOT OF SIGMA X COMPONENT STRESS 
SCALE - 1.0000 PLOT TIME AND DATE - 11:31:06 



'DENOTES HIDDEN 



CJ LN\FNL - Sad ioi Surface 



94/154 



196 




x_ — — H- — 



• • • 


LEGEND •" 




KS1 


A 


3 . 00 


B 


2.00 


C 


1 . 00 


D 


. 00 


E 


-1.00 


F 


• 2.00 





- J. 00 


H 


-4.00 


1 


-5.00 


J 


- 6. 00 


K 


-7.00 


MAX 


3.15 


MIN 


-7.01 



T.TLE NASA 22*. FAN DEFAULT BC S HOT- TO- COLD ( M. CJ LNF. FNL - S.c. i.. S.,r.c. 
CONTOUR PLOT OF SIGMA Y COMPONENT STRESS 
SCALE = 1.0000 PLOT TIME AND DATE = I1:S1:0« 94/154 



197 




• • • 


LEGEND • 




KS I 


A 


3 1.00 


B 


24. 00 


C 


17.00 


D 


10. 00 


E 


3. 00 


F 


-4.00 


C 


-11.00 


H 


- I *. 00 


MAX 


3 1.30 


MIN 


- 20. OS 



TITLE NASA 22 ia FAN DEFAULT BOS HOT- TO- COLD IMiCl LNF.FNL 
CONTOUR PLOT OF SIGMA 2 COMPONENT STRESS 
SCALE - 1.0000 PLOT TIME AND DATE > 11:31:12 94/IJ4 



Suet ioi Stir face 



198 




• « • 


LEGEND • 




KS I 


A 


33.00 


B 


2 ». 00 


C 


23. 00 


D 


18 00 


E 


13.00 


F 


8. 00 


C 


3. 00 


MAX 


33.29 


MIN 


- . 03 



TITLE NASA 22 i. FAN DEFAULT BO S HOT- TO- COLD 1 M. C] 
CONTOUR PLOT OF MAXIMUM PRINCIPAL STRESS 
SCALE = 1.0 00 PLOT TIME AND DATE » 11:31:14 



LNF. FNL - Sncl 



94/154 



Slide: 



199 




• • # 


LEGEND • 




KS I 


A 


11.00 


B 


». 00 


C 


7. 00 


D 


S. 00 


E 


3. 00 


F 


1 . 00 


G 


-1.00 


MAX 


11.04 


MIN 


-2.23 



TITLE NASA 22 i. FAN DEFAULT EC'S HOT- TO- COLD I Mi C) INF. FNL - S.el io. S.lf.c. 
CONTOUR PLOT OF SECOND PRINCIPAL STRESS 
SCALE - 1.0000 PLOT TIME AND DATE - 11:31:16 94/IJ4 



200 




. . . 


LEGEND • 




KS I 


A 


. 00 


B 


-4.00 


C 


-8.00 


D 


- 12.00 


E 


-16.00 


F 


-20.00 


G 


- 24. 00 


MAX 


. 01 


M1N 


-27.0 1 



TITLE NASA 22i. FAN DEFAULT BC'S HOT- TO- COLD I Mt C] LNF. FNL - S.el i.i Surf.ee 
CONTOUR PLOT OF MIN I MUM PR I NC I P AL STRESS 
SCALE = 1.0000 PLOT TIME AND DATE . 11:31:17 94/134 



201 




TITLE NASA 22 it FAN DEFAULT BC S HOT- TO- COLD (M.C) LNF. FNL - Snct ioa Sitdce 
PLOT OF DEFLECTED SHAPE 
SCALE « 1.1000 PLOT TIME AND DATE « 11:31:19 94/154 



202 



TITLE NASA 22" LOW NOISE FAN RIG 
TIME AND DATE 15:49:19 94/157 

GEOSCTRY PLOT SCALE-I.06« 



10 0* Nd 



LOAD SET 




SUBSTRUCTURE 


NAUC 


CODE 


DISK 


A 


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E 


BOLT 


F 



203 



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APPENDIX F 
RESULTS OF DYNAMIC ANALYSIS BLISK 



Figure Al: Mode Shape of Blade Mode 1 






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Figure A2: Mode Shape of Blade Mode 2 






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Figure A3: Mode Shape of Blade Mode 3 






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Figure A4: Mode Shape of Blade Mode 4 



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Figure A5: Mode Shape of Blade Mode 5 



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Figure A6: Mode Shape of Blade Mode 6 






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Figure A7: Mode Shape of Blade Mode 7 






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255 



Figure A8: Mode Shape of Blade Mode 8 






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256 



Figure A9: Mode Shape of Blade Mode 9 






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Figure A 10: Mode Shape of Blade Mode 10 



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Figure A 12: Node Line Plot of Blade Mode 1 - Suction Side 








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Figure A14: Node Line Plot of Blade Mode 2 - Suction Side 



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Figure A15: Node line Plot of Blade Mode 3 - Pressure Side 



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Figure A16: Node Line Plot of Blade Mode 3 - Suction Side 







X 







i~ 










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265 



- Figure A18: Node Line Plot of Blade Mode 4 - Suction Side 



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Figure A20: Node Line Plot of Blade Mode 5 - Suction Side 



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269 



Figure A22: Node Line Plot of Blade Mode 6 - Suction Side 




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271 



Figure A24: Node Line Plot of Blade Mode 7 - Suction Side 








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Figure A26: Node Line Plot of Blade Mode 8 - Suction Side 




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[ Figure A28: Node Line Plot of Blade Mode 9 - Suction Side 




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Figure A29: Node Line Plot of Blade Mode 10 - Pressure Side- 




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r Figure A30: Node Line Plot of Blade Mode 10 - Suction Side 




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Figure A31: Dynamic Stress Plot of Blade Mode 1 - Pressure Side- 



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Figure A32: Dynamic Stress Plot of Blade Mode 1 - Suction Side 



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Figure A34: Dynamic Stress Plot of Blade Mode 2 - Suction Side 



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APPENDIX H 
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Appendix C: Dynamic Stress Contour Plots 



List of Figures 

Cl - C2: Baseline & Aft Vanes - Mode 1 (IB) 
C3 - C4: Baseline & Aft Vanes - Mode 2 (11) 
C5 - C6: Baseline & Aft Vanes - Mode 3 (2B) 
C7 - C8: Baseline & Aft Vanes - Mode 4 (2T) 
C9 - CIO: Baseline & Aft Vanes - Mode 5 (3B) 
Cl 1 - C12: Baseline & Aft Vanes - Mode 6 (3T) 
C13 - C14: Baseline & Aft Vanes - Mode 7 
C15 - C16: Baseline & Aft Vanes - Mode 8 
C17 - C18: Baseline & Aft Vanes - Mode 9 
C19 - C20: Baseline & Aft Vanes - Mode 10 

C21 - C22: Swept Vane - Mode 1 (IB) 
C23-C24: Swept Vane -Mode 2 (IT) 
C25 - C26: Swept Vane - Mode 3 (2B) 
C27 - C28: Swept Vane - Mode 4 (2D 
C29 - C30: Swept Vane - Mode 5 (3B) 
C3 1 - C32: Swept Vane - Mode 6 (3T) 
C33-C34: Swept Vane - Mode 7 
C35-C36: Swept Vane - Mode 8 
C37-C38: Swept Vane - Mode 9 
C39 - C40: Swept Vane - Mode 10 

C41 - C42: Swept & Leaned Vane- Mode 1 (IB) 
C43 - C44: Swept & Leaned Vane- Mode 2 (IT) 
C45 - C46: Swept & Leaned Vane- Mode 3 (2B) 
C47 - C48: Swept & Leaned Vane- Mode 4 (2T) 
C49 - C50: Swept & Leaned Vane- Mode 5 (3B) 
C5 1 - C52: Swept & Leaned Vane- Mode 6 (4B) 
C53 - C54: Swept & Leaned Vane- Mode 7 
C55 - C56: Swept & Leaned Vane- Mode 8 
C57-C58: Swept & Leaned Vane- Mode 9 
C59 - C60: Swept & Leaned Vane- Mode 10 



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